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3 LIBRARY
$100 (a Michigan State
University

This is to certify that the
dissertation entitled

FRICTION ENERGY ABSORPTION IN FIBER REINFORCED
COMPOSITES

presented by

THOMAS JAY BRIMHALL

has been accepted towards fulﬁllment
of the requirements for the

PhD degree in EngineerimMechanlcs

 

 

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Major Professor“! Signature

(792;;{g 05/05.-

Date

MSU Is an Aﬁimatlw Action/Equal Omartunly Institution

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PLACE IN RETURN Box to remove this checkout from your record.
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DATE DUE DATE DUE DATE DUE

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

2105 cmﬁmeoueindd-sz

FRICTION ENERGY ABSORPTION IN FIBER REINFORCED
COMPOSITES

By

Thomas Jay Brimhall

A DISSERTATION

Submitted to
Michigan State University
In partial fulﬁllment of the requirements
for the degree of

DOCTOR OF PHILOSOPHY
Department of Mechanical Engineering

2005

ABSTRACT

FRICTION ENERGY ABSORPTION IN FIBER REINFORCED
COMPOSITES

By

Thomas Jay Brimhall

Energy absorption of ﬁber reinforced composite structures is of interest to the
automotive industry as their speciﬁc energy absorption (SEA), i.e. the energy absorption
capability per unit mass, is higher than many metallic counterparts. However, the SEA of
composite structures has been observed to decrease under dynamic crush loading when
compared with quasi-static compression. This is different from metallic structures. For
example, carbon ﬁber/vinyl ester composite crush tubes crushed at 2.0 m/sec were
observed to have SEA of 23.8 J/gm, a decrease in SEA of 6.6 J/gm or 21.7% compared
with quasi-statically loaded SEA of 30.4 J/gm.

Glass ﬁber/vinyl ester composite crush tubes were investigated with quasi-static
compression and energy-absorbing modes were identiﬁed. The observed energy
absorbing modes included tube comer splitting, composite delamination, matrix damage
due to bending, and sliding friction of the composite with the plug type crush trigger.
These same energy-absorbing modes were observed in quasi-statically compressed and
dynamically crashed carbon/vinyl ester composite crush tubes. Energy absorption
attributable to corner splitting at quasi-static compression was estimated using standard
tensile test results. Corner splitting was estimated to absorb less that 1% of the total

energy absorbed by both the glass ﬁber composite and the carbon ﬁber composite crush

tubes. Energy absorption attributable to delamination was estimated using the mode 11
(shear mode) strain energy release rate obtained using the end notch ﬂexure (ENF) test.
Under quasi-static compression, the glass ﬁber composite delamination SEA was found
to be 1.31 J/gm or 6.4% of the total tube SEA. For carbon ﬁber composite crush tubes,
the delamination SEA was found to be 0.84 J/gm or 2.8%f the total tube SEA.
Experiments seemed to suggest that sliding ﬁiction played an important role in
the energy absorption of composite crush tubes. In an attempt to separate the sliding
friction SEA from the SEA attributable to matrix damage due to bending, an innovative
strip testing ﬁxture was designed, fabricated, and tested and is presented in this thesis.
Carbon ﬁber composite SEA, under quasi-static compression, attributable to matrix
damage due to bending was found to be 18.9 J/gm or 62.2% of the total tube SEA. Under
dynamic crush, the SEA attributable to matrix damage due to bending was found to be
18.6 J/gm or 78.1% of the total tube SEA. Carbon ﬁber composite SEA attributable to
sliding ﬁiction under quasi-static compression was found to be 10.6 J/gm or 34.8% of the
total tube SEA. Under dynamic crush, the SEA attributable to sliding friction was found
to be 4.3 J/gm or 18.1% of the total SEA. The decrease in sliding friction SEA of 6.3
J/gm accounted for nearly all of the decrease in tube SEA of 6.6 J/gm between dynamic
crush and quasi-static compression. Sliding ﬁ'iction was concluded to be responsible for

the decrease in overall tube SEA from quasi-static compression to dynamic crush.

 

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DEDICATION

I dedicate this dissertation to my wife, Carol, and to my children Elizabeth,

Nicole, and Michael. They were the inspiration for my efforts and success.

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ACKNOWLEDGEMENTS

I would like to acknowledge and thank the people who supported, encouraged,
and helped me in the achievement described in this thesis:

To Richard Jeryan, my ﬁrst supervisor at Ford Motor Company, for his
suggestions, constructive criticism and coaching that motivated me to look at all
possibilities.

To my current supervisor, Nripen Saha, for his support and encouragement.

To the managers and directors at Ford Motor Company, Alan Taub, Jim Boland,
Priya Prasad, Paul Killgoar, Charles Wu, and Saeed Barbat for their support and
encouragement to pursue a PhD by making this project a part of my employment goals.

To John Jaranson, for his encouragement and assistance in technical graphics
creation.

To technical fabricators of the Ford Scientiﬁc Research Laboratory shop, Paul
Ostrander and Gary Fleming, for the fabrication of and their design suggestions of the
strip test ﬁxture described in this thesis.

To Mike Starbuck and Rick Battiste of Oak Ridge National Laboratory, National
Transportation Research Center, for conducting dynamic load tests.

To Dan Houston and Ron Cooper of the Ford Scientiﬁc Research Laboratory for
performing quasi-static crush tests.

I would also like to acknowledge and thank the DOE program management team
and the Board and staff of the Automotive Composites Consortium. This work was
sponsored by the Automotive Composites Consortium and the U. S. Department of

Energy, Ofﬁce of Transportation Technologies, Ofﬁce of Advanced Technologies,

030

 

 

 

 

Lightweight Materials Program under Cooperative Agreement number FCOS-

020R22910.

LIST l

LIST ‘

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TABLE OF CONTENTS

LIST OF FIGURES ....................................................................................................................................... X
LIST OF TABLES ..................................................................................................................................... XIII
CHAPTER 1. INTRODUCTION ................................................................................................................. 1
1.1 INTRODUCTION AND MOTIVATION .......................................................................................................... 1
1.2 LITERATURE REVIEW: AUTOMOTIVE CRASHWORTHINESS AND COMPOSITE CRASH TUBES. ................... 8
1.2.1 Composite structures, weight savings, energy absorption. ................................. 8
1.2.2 Design using tubes and SEA. ......................... '. .................................................. 8
1.2.3 SEA of metals and composites .......................................................................... 9
1.2.4 Energy absorption modes of metals and composites. ..................................... 10
1.2.5 Composite tube static vs. dynamic SEA. ........................................................ 13

1.2.6 Strain rate dependence of steel & composite, dynamic & static composite

SEA. ........................................................................................................................... 14
1.2.7 Composite tube energy absorbing modes. ........ 14
1.2.8 SEA of composites and energy absorbing modes, computer simulation. ....... 15

1.3 STATEMENT OF PROBLEM AND OBJECTIVES. ............................................................................ 16
1.4 ORGANIZATION OF DISSERTATION ................................................................................................... 17
CHAPTER 2. DAMAGE OF CRUSH TUBE: MACROSCOPIC AND MICROSCOPIC ....................... 18
2.1 MATERIALS ........................................................................................................................................ 18
2.2.1 Static Crush Tube Tests .................................................................................. 23
2.2.2 Dynamic Carbon Fiber Composite Crush Tube Tests .................................... 26

2.3 MATERIAL TESTING ........................................................................................................................... 31

Vii

 

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2.4 QUASI-STATIC CRUSH TUBE TEST RESULTS ...................................................................................... 38

2.4.1 Introductory Remarks: ...................................................................................... 38
2.4.2 Glass Fiber Composite Quasi-Static Crush Tube Test Results: ........................ 40
2.4.3 Carbon Fiber Composite Crush Tube Test Results ......................................... 43

2.5 MACROSCOPIC DAMAGE MODES ....................................................................................................... 48
2.6 MICROSCOPIC DAMAGE MODES ........................................................................................................ 50
CHAPTER 3. CHARACTERIZATION OF INDIVIDUAL DAMAGE MODES .................................... 55
3.1 INTRODUCTORY REMARKS ................................................................................................................. 55
3.2 CRUSH TUBE TESTING ....................................................................................................................... 56
3.2.1 General SEA Calculation ................................................................................ 56
3.2.2 Crush Triggers ................................................................................................. 57
3.2.3 Crush Tube Energy Absorbing Modes ............................................................ 63

3.3 TAPERED TRIGGER .............................................................................................................................. 72
3.3.1 Introductory Comments ................................................................................... 72
3.3.2 Tapered Trigger Design .................................................................................. 72
3.3.3 Tapered Trigger Crush Tube Results .............................................................. 77

3.4 TAPERED TRIGGER SUMMARY ............................................................................................................ 84
CHAPTER 4. ANALYTICAL SOLUTION TO THE TAPERED TRIGGER CRUSH TEST .................. 86
4.1 INTRODUCTION ................................................................................................................................... 86
4.2 FREE BODY DIAGRAM ......................................................................................................................... 86
CHAPTER 5. FRICTION AND DAMAGE PROCESS .............................................................................. 95
5.1 LUBRICATED CRUSH TUBES, QUASI-STATIC AND DYNAMIC LOADING .......................................... 95
5.2 STRIP TESTING ................................................................................................................................. 102
5.2.1 Strip Test Introduction .................................... 102
5.2.2 Strip Test Specimen and Fixture Description ................................................ 105

viii

5.2.3 Quasi-Static Strip Tests .................................................................................. 108

5.2.4 Dynamic Strip Tests ....................................................................................... 113
5.2.5 Strip Test Results Discussion and Summary ................................................. 121

5.3 ENERGY BALANCE ........................................................................................................................... 124
CHAPTER 6. CONCLUSIONS AND FUTURE RESEARCH ............................................................... 129
6.1 INTRODUCTORY REMARKS ............................................................................................................... 129
6.2 CONCLUSIONS .................................................................................................................................. 130
6.3 FUTURE RESEARCH .......................................................................................................................... 132
REFERENCES ............................................................................................................................................ 133
APPENDIX A ............................................................................................................................................ 136
APPENDIX B ............................................................................................ . ................................................. 142
APPENDIX C ............................................................................................................................................ 153
APPENDIX D ............................................................................................................................................. 160

LIST OF FIGURES

Figure 1.1 Exploded View of generic vehicle showing energy absorbing components. ..... 2
Figure 1.2 Crushed steel and aluminum crush tubes. ......................................................... 4
Figure 1.3 Steel rail crush load-displacement plot ............................................................ 4
Figure 1.4 Average crush load vs. crush speed for glass ﬁber composite. ....................... 7
Figure 2.1 Stitched mat schematic. ................................................................................. 21
Figure 2.2 Tri-axial braid schematic. .............................................................................. 21
Figure 2.3 Standard crush tube triggers. ......................................................................... 24
Figure 2.4 Tube crush test schematic ................ 25
Figure 2.6 TMAC installation at NTRC. ........................................................................ 29
Figure 2.7 Load cell and load washer location on TMAC .............................................. 30
Figure 2.8 TMAC showing bolted crush trigger ............................................................. 32
Figure 2.9 Typical glass ﬁber composite tensile load-displacement curve. .................... 35
Figure 2.10 Typical carbon ﬁber composite 90° tensile load-displacement curve. ........ 37
Figure 2.11 Typical glass ﬁber composite crush tube. ................................................... 39
Figure 2.12 Typical glass ﬁber composite quasi-static crush tube. ................................ 42
Figure 2.13 Typical carbon ﬁber tube quasi-static load-displacement plot.................... 45
Figure 2.14 Typical carbon ﬁber crush tube dynamic load-displacement plot ............... 47

Figure 2.15 Quasi-statically compressed glass ﬁber tube showing delamination, corner
splitting, and matrix damage due to bending. ............................................................ 52
Figure 2.16 Quasi-statically compressed carbon ﬁber tube Showing delamination, comer

splitting and matrix damage due to bending. ............................................................. 53

Figure 2.17 Dynamically compressed carbon ﬁber composite crush tube Showing

delamination, comer splitting and matrix damage due to bending. ........................... 54
Figure 3.1 Glass ﬁber composite tube showing bevel trigger. ....................................... 58
Figure 3.2 Crush tube bevel detail. ................................................................................ 60
Figure 3.3 Glass ﬁber composite tube with bevel and plug trigger. ............................... 61
Figure 3.4 Drawing of standard plug trigger. ................................................................. 62
Figure 3.5 Typical quasi-statically compressed glass ﬁber composite tube. .................. 64
Figure 3.6 Typical quasi-statically compressed carbon ﬁber composite tube. ............... 65
Figure 3.7 Tapered trigger drawing. ............................................................................... 74
Figure 3.7 Tapered trigger. ............................................................................................. 75
Figure 3.8 Compressed glass ﬁber composite tube using tapered trigger. ..................... 76

Figure 3.9 Glass ﬁber composite tubes, load-displacement plot for quasi-statically

loaded with tapered trigger ......................................................................................... 79
Figure 3.10 Crushed carbon ﬁber composite tube using tapered trigger. ....................... 83
Figure 4.1 Tapered trigger tube crush free body diagram. ............................................. 88
Figure 4.2 Section details for tapered trigger .................................................................. 92

Figure 5.1 Carbon ﬁber composite tube crush quasi-static (0.0508 m/min) coated with
zinc stearate. ............................................................................................................... 98

Figure 5.2 Carbon ﬁber composite tube crush dynamic loading coated with zinc stearate.

.................................................................................................................................. 101
Figure 5.3 Strip test ﬁxture schematic. .......................................................................... 104
Figure 5.4 Strip test coupon. .......................................................................................... 106
Figure 5.5 Strip test ﬁxture with coupon. ....................................................................... 106

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Figure 5.6 Typical tested strip coupon ............................................................................ 107
Figure 5.7 Typical load-displacement plot for quasi-static loaded ................................. 112
(0.0508 m/min) carbon composite strip with locked roller D ......................................... 112
Figure 5.8 Typical carbon composite strip load-displacement dynamic loading with 1
locked roller D .......................................................................................................... 115

Figure 5.9 Typical strip load-displacement dynamic loading and free roller D. ............ 120

xii

 

 

 

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LIST OF TABLES

Table 2.1 Composite tube characteristics. ...................................................................... 20
Table 2.2 Crush tube geometry. ...................................................................................... 22
Table 2.3 Glass ﬁber composite plaque mechanical properties. ..................................... 34
Table 2.4 Carbon ﬁber plaque composite mechanical properties. .................................. 36

Table 2.5 Glass ﬁber composite crush tube SEA, quasi-static loading rate (0.0508
m/min). ....................................................................................................................... 41

Table 2.6 Carbon ﬁber composite crush tube SEA, quasi-static loading rate (0.0508

m/min). ....................................................................................................................... 44
Table 2.7 Carbon ﬁber composite crush tube SEA dynamic loading. ............................ 46
Table 3.1 Glass ﬁber composite corner splitting SEA ..................................................... 67
Table 3.2 Carbon ﬁber composite corner splitting SEA. ................................................. 67
Table 3.3 Glass ﬁber composite SEA due to delamination. ........................................... 71
Table 3.6 Carbon ﬁber composite SEA due to delamination. ........................................ 71

Table 3.7 Glass ﬁber composite crush tube SEA using tapered trigger quasi-static load
rate (0.0508 m/min). ................................................................................................... 80

Table 3.8 Carbon ﬁber composite crush tube SEA using tapered trigger quasi-static

loading rate (0.0508 m/min) ....................................................................................... 82
Table 3.9 Energy absorption summary for quasi-static loading. .................................... 85
Table 4.1 Variable values. .............................................................................................. 93

Table 5.1 Carbon composite crush tube SEA with zinc stearate coating at quasi-static

loading (0.0508 m/min) .............................................................................................. 96

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Table 5.2 Carbon ﬁber composite crush tube SEA with zinc stearate low friction coating,
dynamic loading. ...................................................................................................... 100
Table 5.3 Carbon tube test summary non—lubricated VS. lubricated with zinc stearate. 100
Table 5.4 Carbon ﬁber composite strip SEA, locked roller D quasi-static loading (0.0508
m/min). ..................................................................................................................... 109

Table 5.5 Carbon ﬁber composite strip SEA, locked and lubricated roller D, quasi-static

loading (0.0508 m/min) ............................................................................................ 110
Table 5.6 Carbon ﬁber composite strip SEA, free roller D, ........................................... 111
quasi-static loading (0.0508 m/min). .............................................................................. 111

Table 5.7 Carbon ﬁber composite strip SEA, locked roller D, dynamic loading (2.0
m/sec). ...................................................................................................................... 114
Table 5.8 Carbon ﬁber composite strip SEA, locked roller D, dynamic loading (0.5
III/sec). ...................................................................................................................... 1 14
Table 5.9 Carbon ﬁber composite strip SEA, locked and lubricated roller D, dynamic
loading (2.0 m/sec). .................................................................................................. 117
Table 5.10 Carbon ﬁber composite strip SEA, free roller D, dynamic loading (2.0
m/sec). ...................................................................................................................... 1 19
Table 5.11 Carbon ﬁber composite strip SEA, free roller D, dynamic loading (0.5 m/sec).
.................................................................................................................................. 1 1 9
Table 5.12 Strip test results summary. ........................................................................... 123
Table 5.13 Carbon composite strip and crush tube energy balance quasi-static loading.

.................................................................................................................................. 127

xiv

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Table 5.14 Carbon composite strip and crush tube energy balance dynamic loading

(2.0 m/sec). ............................................................................................................... 128
Table 5.15 Quasi-static vs. dynamic loaded carbon composite tube crush dynamic load

rate (2.0 m/sec). ........................................................................................................ 128

XV

 

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1.1

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CHAPTER 1. Introduction

1.1 Introduction and motivation

The body structure of the modern automobile is required to protect the occupants
in the event of a crash. To this end, automobiles are designed with an occupant cell that
serves as protection for the passengers in the event of a crash. Within the passenger
capsule, interior components are designed to protect passengers by absorbing energy due
to impact with passengers. Such items include a collapsible steering column, padding on
the instrument panel, and deployable airbags. Seat belts restrain passengers in the event
of a collision preventing contact with solid surfaces such as the instrument panel and
windshield.

Front structures are designed to absorb crash energy by progressively collapsing
during the crash event. The energy absorbing front structure consists of upper and lower
rails, front crossmember, aprons, and secondary body add-on parts (Figure 1.1). The
primary energy absorption structures are the upper and lower rails. The front
crossmernber ties the left and right lower rails providing increased torsion stifﬁress. The
ﬁ'ont aprons provide additional torsion stiffness by linking the upper and lower rails
together. The front apron also absorbs crash energy during the crash event. Bumpers are
designed to transfer load ﬁom the point Of impact to the lower front rails. Bumpers
provide protection to the automobile from low speed impacts and absorb small amounts
of energy in hi gh-speed frontal crashes.

Experience has shown that simple structures such as tubes can be used to help
characterize the behavior Of more complex structures. The primary measurement Of the

energy absorption capability of a structural material is speciﬁc energy absorption (SEA).

 

 

 

CROSSMEMBER

Figure 1.1 Exploded View Of generic vehicle showing energy absorbing components.

The SEA is deﬁned as the quantity of energy absorbed divided by the mass of the crash
damage or deformed material. Crush tubes made from the same materials as the front
rails are used for experimental measurement of energy absorption. A simple tube is used
to characterize energy absorption because of its similarity in conﬁguration with
automotive energy absorbing structures such as the front rails. Typical post-test crush
tubes made from steel and aluminum are presented in Figure 1.2. Metals used in
automotive energy absorbing structures absorb energy primarily by plastic deformation.
Note that there is extensive bending with no ﬁ'acture of the metal observed. A load-
displacement plot for two automotive ﬁont rail tests is presented in Figure 1.3. In
addition, the dynamic behavior of steel structures is routinely predicted by computer
analyses.

As an alternative to steel, composites are being considered more frequently for
structural applications in automobiles because of their lower mass, high speciﬁc stiffness,
tailorable mechanical properties, improved noise vibration and harshness (NVH) and
simpliﬁed assembly due to a smaller number‘of parts due to parts consolidation. In fact,
the chassis of many racecars such as Formula 1 and Indy cars are primarily fabricated
from carbon ﬁber composite. Composites have not yet come into wide use in every day
automotive structural applications because Of high cost, slow fabrication cycle times,
Slow fastening methods (adhesives and mechanical fasteners), recyclability issues, and
the inability to predict dynamic crash behavior.

Before a design for an automobile is approved, conformance to numerous
requirements must be demonstrated by physical testing. The requirement of interest in

this work is crash energy absorption by the front rails. A sure method to ensure

 

Figure 1.2 Crushed steel and aluminum crush tubes.

Crash Force (Ibo

 

0 2 4 6 8 10 12 14
Displacement (in)

Figure 1.3 Steel rail crush load-displacement plot.

conformance to requirements is to crash test a variety Of rail designs until satisfactory
energy absorption is achieved. The disadvantage to this approach is the high cost and
long time needed for thorough investigation. A more systematic approach for
preliminary analysis is to perform computer analysis on a variety of materials and front
rail architectures to identify the best candidates. Before structures can be analyzed, it is
necessary that characteristics, such as SBA and crush load level, be measured. This
cannot be accomplished using complex shaped structures found in the front automobile
structure. It is simpler to analyze and test a simple tube of uniform cross section that is
similar to that of the ﬁ'ont rail. Such a simple tube is less costly to fabricate and allows
investigation of a variety of materials while minimizing geometry effects. This approach
has been used with success with metals and with mixed results using composites.

In contrast with metals, composites typically do not exhibit signiﬁcant plastic
deformation. Both the composite reinforcement and polymeric matrix typically exhibit
elastic stress-strain behavior up to fracture. Composites, by their nature, have multiple
damage modes. Any of these modes has the potential for energy absorption.
Characterization Of composites to be used as energy absorbing structures has been
attempted using the techniques used to characterize steel structures. Speciﬁcally, simple
tubes of a speciﬁc composite architecture are tested statically and dynamically using the
same procedures and equipment used for steel crush tubes.

The strain-rate effect observed in most steels used in automotive structures causes
the average crush load to increase with increasing crush speed. Metal structures have
been well characterized and experimental data have been successfully used with

computer aided engineering (CAB) software to predict energy absorption of complete

 

 

C03

 

automotive structures. Also, strain-rate dependence of metallic materials has been
accounted for in many non-linear CAE ﬁnite element analysis (F EA) solvers such as
RADIOSS® and LSDYNA®. SEA characterization of metal structures begins with
coupon testing in static and dynamic crush testing. The results Of these tests are used for
calibration of CAE models of the crush tube test. Larger more complex sub-component
tests and analyses are performed leading to full vehicle CAE analyses and crash testing.
Similar analysis success has not been demonstrated with composite materials in spite of
numerous attempts to develop CAE software that can accurately model the dynamic
crush behavior of composite materials.

Experiment has shown that the SEA of a composite structure when tested
statically can be signiﬁcantly different than observed in dynamic experiments. In most
cases, the dynamic SEA is much less than the static SEA for composite materials. A plot
of average crush load vs. crush speed for two ﬁberglass ﬁber composites is presented in
Figure 1.4, Ref. (1). AS part of the development of analysis of energy absorption of
composite structures, it is necessary to identify and measure dynamic modes of energy
absorption. Differences in SEA between dynamic and static testing must be explained

and a method developed for each material system to predict energy absorption behavior.

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

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Figure 1.4 Average crush load vs. crush speed for glass ﬁber composite.

1.2 Literature review: Automotive crashworthiness and composite crash tubes.

1.2.1 Composite structures, weight savings, energy absorption.

Composite structures have been investigated for use in energy absorbing
structures with the goal of reducing weight, simplifying assembly, and yet meeting
requirements for crash energy absorption. Johnson Ref. (2) reported that polymeric
composites used for large automotive structures Offer advantages in weight reduction and
part consolidation. Speciﬁcally, the Automotive Composite Consortium (ACC) Focal
Project 1 Ford Escort front-end structure made from RTM (resin transfer molded) glass
reinforced polymeric composite had a 25% weight improvement compared with the
baseline Ford Escort steel structure. In addition, there was a reduction in the number of
parts from 90 to 2. Botkin et a1. Ref. (3) reported that ACC Focal Project 1 demonstrated
that the ﬁ'ont-end structure could absorb the necessary energy during crash. This project
by the ACC demonstrated that a vehicle with a composite production feasible ﬁont
structure could meet crash requirements with a reduction in weight and number of parts.
1.2.2 Design using tubes and SEA.

Efforts have been made to improve the efﬁciency of vehicle development by
testing components and coupons rather than full vehicle structures. Browne et al. Ref. (4)
stated that producing a variety of automotive front-end structures and crashing them
under different conditions is time consuming. A convenient testing approach is to use
simple crush tubes that are crushed using a plug initiator under static and dynamic
conditions. The goal is to extrapolate the results of simpler tests to full-scale
automobiles. It is reasoned that SEA, similar to elastic modulus, strength, density, etc., is

a material property and crushing any well-deﬁned structure under well-controlled

conditions will provide the SEA for a speciﬁc composite. Once the SEA is measured, the
result could be applied to the design of full size structures.
1.2.3 SEA of metals and composites.

Measurement Of SEA is routinely performed on metal and composite crush tubes.
Composites crush tubes have been shown to have SEA that is ﬁ'equently higher than steel
and aluminum crush tubes. Derek Hull Ref. (5) reported that the SEA for mild steel
crush tubes was approximately 28.5 J/gm and 22.5 J/gm for 50 mm cylinders with
thicknesses of 1.0 to 2.0 mm. The Automotive Aluminum Crash Energy Management
Manual Ref. (6) reported the SEA for a variety Of aluminum tube architectures and
aluminum alloys ranged from 16 J/gm to 29.5 J/gm. Hull also reported SEA for glass
cloth-epoxy at 64 J/gm, 73.5 J/gm for chopped glass-polyester, 80 J/ gm for oriented
glass-polyester, and 130 J/ gm for carbon ﬁber-epoxy. Hull also reported that composite
crush tube geometry inﬂuences SEA in contrast with steel structures where the SEA
seems to be independent of geometry. The composite specimens tested were similar to
the steel and aluminum test tubes, 50 mm cylinder with wall thiclmess 1.0 to 2.0 mm.
Hull Ref. (5) also stated that composites offer the potential for energy absorption
structures with reduced weight based on the higher SEA observed for a range of
composites. Frequently, damage modes, such as structural instability, which reduces
energy absorption, are observed. For composites to absorb the maximum crash energy,
they must crush progressively and in a controlled manner such as Observed in well-

designed metal structures.

1.2.4 Energy absorption modes of metals and composites.

Energy absorption by metals is primarily due to plastic deformation. The stress-strain
plot for steel used for automotive structures typically has a plateau region corresponding
to nearly constant loading with increasing strain. The region under this plot represents
energy absorbed during progressive crush. When composites are tested as coupons, the
stress-strain plot shows elastic behavior with little or no plastic deformation prior to
fracture. Yet, when a plot of load-displacement for a composite tube (author generated)
is examined, Figure 1.5, a plateau region Of nearly constant load occurs that appears
similar to that of a steel tube, Figure 1.2. Thornton Ref. (7) described composite material
energy absorbing mechanisms and estimated the relative amount of energy absorption
attributable to each of these mechanisms: the creation of fracture surfaces less than 1%,
friction 25% to 50%, and the balance Of energy absorbed due to plastic deformation of
the matrix. Thornton also postulated that Speciﬁc energy absorption was not affected by
crush rate as long as damage modes remained the same. Farley et al. Ref. (8), in contrast
to Thornton, stated that the coefﬁcient of ﬁiction between the composite and the crushing
surface is a function of crushing speed. This suggests that the effect Of crush speed on
SEA is dependent on the composite material properties. Farley stated that a friction
related energy absorption mechanism is due to the relative motion between adjacent
lamina bundles that slide against each other. Farley also stated that the magnitude of
energy absorption attributable to these friction mechanisms had not been quantiﬁed. In a
separate paper, Farley et al. Ref. (9) stated that the primary energy absorption
mechanisms are fracturing of lamina bundles and crack growth and that friction is

considered a secondary energy absorption mechanism. Mamalis et al. Ref. (10) estimated

10

friction energy absorption to be 50% of the energy absorbed. Mamalis found that the
dynamic coefﬁcient of friction to be higher than the static coefﬁcient of ﬁiction. In a
separate paper, Mamilis et al. Ref. (11) found energy absorption due to ﬁiction was from
frictional resistance to axial sliding between adjacent laminae, frictional resistance to
penetration of the annular debris wedge formed when a Circular tube is crushed against a
ﬂat platen and frictional resistance to ﬁonds sliding across the platen. Mamilis postulated
‘ that energy absorption due to friction mechanisms was the most signiﬁcant of the various

energy absorption mechanisms. It is apparent that a method to quantify energy

absorption due to ﬁiction is needed.

11

bond 0“)

20.000

1 5.000

1 0.000

5,0!”

Tube 4278-2-11

 

 

 

 

 

 

MVMV “WWII TEA/MAN V

 

 

J

 

 

 

 

 

 

T

l

 

 

0 so 100 150
Dlsplacenclt (rum)

Figure 1.5 Typical load-displacement plot for carbon ﬁber crush tube.

12

200

250

 

1.2.5 Composite tube static vs. dynamic SEA.

Boeman and Caliskan Ref. (1) showed that the static crush load was
approximately twice the dynamic load for crush speeds ranging ﬁ'om 2 m/sec to 7m/sec.
This data is presented in Figure 1.4. Chadwick and Caliskan Ref. (12) observed that
crushed vinyl-ester glass composite tubes absorbed energy by matrix cracking, ﬁber
fracture, interfacial debonding, ﬁber pullout, and friction. Chadwick and Caliskan stated
that analysis to determine which energy absorption mode absorbed more energy was
impossible. Dynamic crushing was accomplished using a drop tower with an impact
velocity of 4.99 m/s. Static testing was performed using a standard materials testing
machine at 0.0508 m/min. The dynamic energy absorption was found to be signiﬁcantly
less than static energy absorption. The crush load was found to actually increase as the
crush velocity decreases indicating energy absorption is indeed higher at lower velocities.
Chadwick and Caliskan concluded that the difference in SEA between static and dynamic
conditions was real and not a testing artifact. Johnson et al. Ref. (13) reported that the
ratio between static and dynamic SEA varied from 0.64 to 1.70 depending on ﬁber
architecture and tube wall thickness. However, the two ﬁber architectures that showed
greater average dynamic SEA than average static SEA also had such a large experimental
coefﬁcient of variation that no conclusions could be drawn as to whether the static SEA
was indeed less than the dynamic SEA. Crush morphology did not vary signiﬁcantly
between static and dynamic testing. This observation, in the light of Thornton's Ref. (7 )
conclusion that SEA would not vary between static and dynamic loading as long as

damage modes were the same, would suggest that SEA should remain constant and not

13

vary as was observed by Johnson Ref. (13). Johnson‘s observations Show that SEA is
affected by crush speed and ﬁber architecture as well as tube thickness.
1.2.6 Strain rate dependence of steel & composite, dynamic & static composite SEA.

Bruce et al. Ref. (14) reported that a variety of automotive steels show strain rate
dependence. Speciﬁcally, the tensile strength was measured to increase by 25 MPa to 50
NIPa per magnitude increase in the strain rate. Weeks and Sun Ref. (15) reported that
AS4/PEEK carbon ﬁber reinforced composite and an increase in tensile strength of 25
MPa per magnitude increase in the strain rate for [i30] laminate and no signiﬁcant strain
rate effect for unidirectional material tested in the ﬁber direction. Todo et a1. Ref. (16)
observed tensile strength increase of approximately 20 MPa per order of magnitude
increase in the strain rate for both carbon ﬁber and glass ﬁber woven composites. These
observations differ with the observed decrease in SEA between dynamic and static SEA
reported by Boeman and Caliskan Ref. (1). A number of sources, as documented above,
Show that reduced dynamic SEA compared to static SEA is contrary to the observed
unchanged or increase in strength with increase in strain rate.

1.2.7 Composite tube energy absorbing modes.

Johnson et al. Ref. (17, 18) identiﬁed energy dissipation modes as ﬁber tearing,
resin fracture, inter-ply delamination and friction against the initiator for structural
reaction injected molded (SRIM) glass tubes crushed under static and dynamic loading.
Browne et a1. Ref. (19) reported that energy dissipation in resin transfer molded (RTM)
crush tubes loaded statically and dynamically was due to resin fracture, ﬁber fracture and
pullout at the four tube corners, resin and resin-ﬁber interface fracture and inter-ply

delamination and ﬁiction against the initiator. Browne also Observed that among crush

14

tubes of different glass ﬁber architectures, the SEA was not proportional to volume
fraction of glass in the composite. This suggests that ﬁber reinforcement fracture is not
the primary energy absorbing mechanism. This also suggests that SEA may not be only a
material property. Clearly, other researchers have Observed multiple modes of energy
absorption of crush tubes tested statically and dynamically without measuring the relative
amount of energy absorption attributable to each energy-absorbing mode. A deeper
understanding of energy absorption modes is necessary before the performance of
designs of energy absorbing automotive structure designs are reliably predicted. "What
are the energy absorbing modes and how much energy is absorbed by each mode?" is a
question that must be answered for each material system and energy absorbing structure.
TO date, no one has reported a method to measure the SEA attributable to each energy-
absorbing mode.

1.2.8 SEA of composites and energy absorbing modes, computer simulation.

Agaram et a1. Ref. (20) reported using a continuum damage scheme to mimic the
effect of damage as a localized stifﬁress reduction. This approach did not result in
satisfactory prediction of energy absorption and requires the user to guess the constitutive
behavior of the damage material. Agaram recommended that predictive computer codes
be developed in conjunction with fundamental studies. The conclusion can be drawn that
computer simulation and prediction of energy absorption due to composite crush cannot
be accomplished until the fundamental energy absorption modes are identiﬁed, described,

and quantiﬁed.

15

1.3 Statement of problem and objectives.

Prediction of dynamic and static energy absorption by composite automotive
structures requires that all energy absorption modes be identiﬁed and quantiﬁed. The
most expedient method of characterizing energy absorbing materials is to use relatively
Simple test coupons that are similar to actual structures so that a transfer function can be
established to correlate testing of test structures with automotive structure designs.

The ﬁrst part of this problem is addressed by the careful Observation of tubes
subjected to crush loading. Energy absorbing modes can be identiﬁed using observation.
The second part of the problem, quantiﬁcation, is more difﬁcult because of the
interaction between energy absorbing modes.

After energy absorption modes are identiﬁed, a test must be designed that will
isolate the modes or groups of modes. It is hypothesized that the energy absorption due
to ﬁiction mechanisms is one of the primary modes of energy absorption for composite
materials. Further, the difference in ﬁiction energy absorption under static loading VS.
dynamic loading is the primary reason that the SEA of a composite varies depending on
loading speed.

The goal of the research presented is to identify and then measure the energy
absorption of each identiﬁable, signiﬁcant mode for both static and dynamic loading.
The measurement of energy absorption is performed by conducting tests on simple test
specimens. These simple tests are designed to isolate single or a limited number of
energy absorptionmodes. The tests are performed under both static and dynamic
loading. The energy absorption from each mode can be summed and compared with the

total energy absorption Observed in crush tests of a component structure with the same

16

composite materials, ﬁber architecture, and crush speed. The relative amount of energy
absorbed due to friction can be compared between dynamic and static tests to
demonstrate if and to what degree ﬁiction plays a part in energy absorption of a
composite automotive structure.
1.4 Organization of Dissertation

The research presented within this thesis is divided into Six chapters. The ﬁrst
chapter provides background information, literature search, and organization. The
second chapter will describe damage of experimental crush tubes of two different
composite systems. Within chapter 2 there will be experimental results of the full crush
tube testing and of the individual simple tests that isolate energy absorption modes.
Chapter 3 will characterize individual energy absorbing modes. Chapter 4 will show an
analytical solution to the tapered trigger test. Chapter 5 will describe the role of ﬁiction
in the crush tube damage and energy absorbing process. Chapter 6 will present

conclusions and recommendations for future study and research.

17

CHAPTER 2. Damage of crush tube: Macroscopic and Microscopic

2.1 Materials

The two composite materials used in this study were ﬁberglass/vinyl ester resin
and carbon ﬁber/vinyl ester. The glass ﬁber reinforcement was in the form of stitched
mat. A schematic of the stitched mat is presented in Figure 2.1. Each ply of stitched
unidirectional ﬁber tow glass mat has three layers of unidirectional glass ﬁber tows held
together with stitching. The stitching holds the dry ﬁber tows together allowing handling
prior to molding. The ﬁber tow direction of each layer is perpendicular to the adjacent
layer. In this case, there are twice as many ﬁbers oriented in the 0° direction as the 90°
direction with the 0° direction parallel the longitudinal direction of the crush tubes. This
conﬁguration is designated in Table 2.1 as [0/90/0]. Six of these stitched plies Were
stacked together by wrapping the stitched glass ﬁber mat around the tube mold mandrel
six times providing the desired ﬁber reinforcement build-up. The wrapping starts and
ends on one Side of the mandrel, resulting in a seam line. The tube was then molded
using the resin transfer molding (RTM) process for the glass ﬁber composite crush tube
used in this study.

Details of the glass ﬁber and carbon ﬁber crush tube geometry are presented in
Table 2.2. All test tubes were manufactured with the RTM process at Excel Pattern Co.
(Detroit, MI). In the RTM process, dry ﬁber reinforcement is placed in a mold which is
then closed. Liquid resin is then injected into the heated, closed mold and allowed to
cure. After cure, the molded part is removed from the mold, trimmed of resin flash, and

machined to ﬁnal length.

18

The carbon ﬁber reinforcement is formed by a machine that braids carbon ﬁber
tows into a ply with axial tows oriented parallel with the longitudinal axis of the crush
tubes and angle plies oriented i45° from the longitudinal axis of the crush tubes. One
feature of the braiding process is that there were no seams in the plies. A schematic of
this reinforcement architecture is presented in Figure 2.2. Two such plies were used to
fabricate the carbon ﬁber composite crush tubes used in this study. The braiding process
is being considered because it is a high-speed method of ﬁber architecture manufacture
and is easily applied to tubular structures with little scrap. Like the glass ﬁber composite,
the carbon ﬁber crush composite tubes were manufactured using the RTM process at

Excel Pattern (Detroit, MI).

19

Table 2.1 Composite tube characteristics.

 

 

 

 

Reinforcement Matrix Architecture Fiber Fiber Laminate
Content Volume Density
(wt. We (%) morn’)
Glass Fiber Ashland [0/90/0]6 65.9 46.4 1.800
Knytex/Hexcel Q-6050 Stitched Mat
Vinyl
Ester
Carbon Fiber Ashland Tri-axial Braid 54.5 43.0 1.419
Fortaﬁ1503 80k Hetron [0/145];
Axial 922 Vinyl
Graﬁl 12k Off-Axis Ester

 

 

 

 

 

 

20

 

 

Figure 2.1 Stitched mat schematic.

 

Figure 2.2 Tri-axial braid schematic.

21

Table 2.2 Crush tube geometry.

 

 

 

 

Tube Outside Tube Nominal Cross
Width and Length Thickness Section
Depth (mm) (mm) (mm) Area
(cmz)
Glass Fiber 55.5 229 4.7 10.1
Knytex/Hexcel
Carbon Fiber 55.5 356 2.4 4.74
Fortaﬁ1503 80k
Axial
Graﬁl 12k

Off-Axis

 

 

 

 

 

22

 

2.2 Test Methods and Facilities
2.2.1 Static Crush Tube Tests

All static crush tube testing was performed at the Ford Scientiﬁc Research
Laboratory using an MTS 810 (servo hydraulic) material test system at ambient room
conditions. The MTS load frame has a 222 KN (50,000 lb.) capacity and uses a 22.2 KN
(5,000 lb.) load cell. All specimen testing and data was controlled, recorded and
analyzed through the use of the SinTech software package, Test Works, on a Dell
personal computer. The crosshead speed was 0.0508 m/min (2 inches/min).

Five glass ﬁber composite tubes were tested with the standard plug type trigger
with a ﬁllet radius of 6 mm (Figure 2.3). The trigger for each composite tube, glass ﬁber
and carbon ﬁber, was, machined to ﬁt inside the tube with a tolerance of 1 mm or less.
The glass ﬁber composite tubes were of square cross section with rounded corners
(outside radius 12 mm). The upper end Of the tube was contacted directly by the upper
platen and was not constrained ﬁ'om movement transverse to the axis of the crush tube.
The plug trigger was placed on the lower platen and was also unconstrained from
transverse displacement. A schematic of the tube crush testing is presented in Figure 2.4.
In all cases, no transverse displacement of either end of the crush tube was observed. NO

dynamic tests were conducted with the glass ﬁber composite crush tubes.

23

 

Figure 2.3 Standard crush tube triggers.

24

 

 

V
@UPPER PLATEN \I

 

CRUSH TUBE

 

 

 

   

STANDARD TRIGGER

 

FIXED LOWER PLATEN

 

Figure 2.4 Tube crush test schematic.

25

2.2.2 Dynamic Carbon Fiber Composite Crush Tube Tests

Dynamic testing was performed using the Test Machine for Automotive
Crashworthiness (TMAC) located at the National Transportation Research Center, Oak
Ridge National Laboratory, Tennessee. The engineer responsible for its operation, Dr. J.
Michael Starbuck of Oak Ridge National Laboratory, wrote the description of the TMAC
presented below.

Typically standard test machines are employed for experiments at quasi-static
rates whereas drop towers or impact sleds are the convention for dynamic rates. These
two approaches bound a regime within which data, for experiments at constant impact
velocity, are not available by conventional experimental practice. This regime is termed
in this thesis the intermediate-rate regime and is deﬁned by impact velocities ranging
from 1 m/s to 5 m/s. Investigation of rate effects within this regime requires experimental
equipment that can supply a large force with constant velocity within these rates. Using a
drop tower or sled at intermediate rates, although technically possible, is problematic due
to the prohibitively large mass required to maintain constant velocity during the crush.
Consequently, the Oak Ridge National Laboratory (ORNL) and the Automotive
Composites Consortium (ACC) collaborated to deﬁne speciﬁcations for a unique experi-
mental apparatus that mitigates the shortcomings of existing equipment. MTS Systems
Corporation designed and built the servo-hydraulic test machine, referred to as the
TMAC. As shown in Figure 2.5, TMAC is uniquely capable of conducting controlled
progressive crush tests at constant velocity in the intermediate velocity range (i.e., less

than 5m/s) owing to the large energy available at those rates and to the sophisticated

26

simulation and control software that permits velocity uniformity to within 10%. The
speciﬁcations for the machine, presented in Figure 2.6, are as follows:

0 No.10ad velocity of 8 m/s constant to within 10% over 115 mm.

o Sustained crush force of 133 KN at 6 m/s constant to within 10% over 115 mm.
0 Sustained crush force of 267 KN at 4 m/s constant to within 10% over 115 mm.

A LabView program was developed to obtain the test data ﬁ'0m two hi gh-speed data
acquisition (DAQ) modules—a 12-bit synchronous card with four channels, and a 16-bit
asynchronous card with eight channels. The program synchronizes the data and is
triggered from the TMAC control software that initiates the impact. One card has higher
signal resolution at the expense of temporal resolution and synchronicity, while the other
card sacriﬁces signal resolution for fast, synchronous measurements. The merits of the
two cards are being assessed.

The functionality of the data system was demonstrated and will be reﬁned and
updated as ﬁrture test requirements are deﬁned, including the capability for recording
strain gage signals. The current DAQ conﬁguration reads one LVDT for displacement, a
load cell and load washer for force, and two accelerometers. The accelerometers were

chosen to provide 1 g/V sensitivity and 10 W sensitivity. 1 g/V is equivalent to 9.81
meter/secZ/Volt. For this study, accelerometer data was not used. The location of the

load cell and load washer is presented in Figure 2.7.

27

 

200 I

/ 1 II", I
L

’ 'r-IA.
,1 .1 r
l

150— *

/ . .
, ,v‘ , ,/ 2'
' l 1 1' 1 /

AA;\\ 1‘\
»~ .q<\
V \“ .\\ -\
g 100- \
w ‘i.
C
LIJ

50.. Tower ,:‘j§777|rﬁbact$i
? TMAC i 1‘
, .

We.

   
  

0 4 8 1 2 1 6

VELOCITY (m/s)

 

 

 

Figure 2.5 Energy plot indicating TMAC’s unique capability of supplying enough energy at the

intermediate rates for controlled, constant-velocity crush.

28

 

Figure 2.6 TMAC installation at NTRC.

29

 

 

114m

 

343m

 

Figure 2.7 Load cell and load washer location on TMAC.

30

The plug trigger used in the crush testing was ﬁxed to the TMAC platen as a
safety precaution per National Transportation Research Center requirements. A photo of
bolted plug initiator is presented in Figure 2.8.

2.3 Material Testing

Material property tests were conducted on both the glass ﬁber composite and
carbon ﬁber composite to characterize the selected test materials. Plaques were
fabricated at the same time as the crush tubes and were used for the materials
characterization. Plaque characterization testing was performed by Delsen Testing
Laboratories Glendale, California. Glass ﬁber composite characterization test results are
presented in Ref. (21) and summarized in Table 2.3. A typical glass ﬁber composite
tensile test in the 90° direction is presented in Figure 2.9. Note that the stress VS. strain
curve is nearly linear over the full range of stress. An essentially linear curve is observed
up to the point of material fracture indicating little energy absorption due to plastic
deformation. Fiber and resin weight content for both glass ﬁber and carbon ﬁber
composites were determined using method ASTM D2584, Ref. (22). This method is used
to determine weight content of resin by burning away the resin leaving only the ﬁber
reinforcement residue. The ﬁber residue is weighed and divided by initial composite
sample weight to determine the ﬁber weight fraction. This method presumes the
presence of only ﬁber and matrix. Fiber sizing and void content are not determined using
this method.

Carbon ﬁber composite characterization test results are presented in Ref. (23) and
summarized in Table 2.4. A typical carbon ﬁber tensile load-displacement plot in the 90°

direction is presented in Figure 2.10. As with the glass ﬁber composite, the curve is

31

 

Figure 2.8 TMAC showing bolted crush trigger.

32

 

 

nearly

plasti

 

 

nearly linear up to material fracture indicating that there is little energy absorption due to

plastic deformation.

33

 

Table 2.3 Glass ﬁber composite plaque mechanical properties.

 

 

 

 

 

 

 

 

 

 

 

 

 

Direction Stress, Max Modulus Poisson's Strain, Fracture
(MPa) (GPa) Ratio Max (%) Energy (J/cm3)
Tensile
0° 492 28.3 0.176 2.41 3.60
90° 189 17.1 0.107 1.34 0.791
Compressive
0° 501 27.2 1.86 2.62
90° 203 18.0 1.15 0.669
Shear
64.0 4.46
Strain Energy jS Gnc
Release Rate (J /cm2) (J /cm2)
0.146 0.123
Physical Density Fiber Resin
Properties (gm/cm3) Weight Weight
Content (‘VQ Content C/o)
1.8002 65.9 32.1

 

 

 

 

 

 

34

 

Tube 3 Secimen 3 (90 deg)
40 . . . - . - . . .

 

Stress (K81)
N
o

 

 

_..A_ A ‘ I L

 

 

O . 1.00
Strain (96)

Figure 2.9 Typical glass ﬁber composite tensile load-displacement curve.

35

2.00

Table 2.4 Carbon ﬁber plaque composite mechanical properties.

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

Direction Stress, Max Modulus Poisson's Strain, Max Fracture
(MPa) (GPa) Ratio (%) Energsy
(J/cm )
Tensile
0° 654 60.33 0.562 1.08 3.623
90° 65.2 8.756 0.109 1.20 0.475
Compressive
0° 376 59.71 -- 0.66 1.289
90° 91.7 10.06 -- 0.89 0.404
Short Beam
Shear
43.1
Strain Energy G"c (J/cmz)
Release Rate
0.1 59
Physical Density Fiber Resin
Properties (gm/cm3) Weight Weight
Content (%) Content (%)
Fiber 1.80
Resin 1.14
Composite 1 .42 54.74 45.26

 

 

 

 

 

 

36

 

Load (lb)

90Tl l

 

1,800

1,600 / ~4~————————4

1
1,400 /

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

0 0.05 0.1 0.15 0.2 0.25 0.3

Displacement (In)

Figure 2.10 Typical carbon ﬁber composite 90° tensile load-displacement curve.

37

2.4 Quasi-Static Crush Tube Test Results
2.4.1 Introductory Remarks:

Design of composite structures requires knowledge of the physical properties of
the composite material. This knowledge is obtained by experimental material
characterization of small coupons to determine mechanical properties such as elastic
modulus, Poisson's ratio, density, strength (tensile, compressive, shear), etc. Changing
the ﬁber material and architecture, resin, relative ﬁber content, and manufacturing
method can vary each of these properties. Typically, test coupons are fabricated ﬁ'om ﬂat
plaques because test coupon fabrication cost is low and the tests are simple to perform.
Coupon testing is usually performed at quasi-static loading rates. Dynamic testing is less
frequently performed on test coupons. However, dynamic testing is routinely performed
on full- and sub-structures. A full structure test would be a crash test of an entire
automobile or major portion of an automobile such as a frame. A sub-structure test is
performed on small part of a full structure such as the front portion of an automobile
frame. Small test structures, usually tubes such as those shown in Figure 2.11, are tested
both statically and dynamically to determine the SEA of the material at a given crush
speed. Tubes are fabricated in various cross sections such as circular, hexagonal,
octangular, rectangular, or square. The primary reason for using rectangular tubes to
characterize SEA is for convenience and Similarity with typical automotive energy
absorbing structures, which are tubular and rectangular. A simple rectangular tube is
more Similar to the front rail used in the crash energy absorbing front structure of an
automobile than a ﬂat coupon. The selection of square crush tubes for this study was

because of their availability and the existence of previous experimental data.

38

 

Ical glass ﬁber composite crush tube.

Figure 2.11 Typ

39

2.4.2 Glass Fiber Composite Quasi-Static Crush Tube Test Results:

Test results for the glass ﬁber tubes are presented in Table 2.5. A typical plot of
load-diSplacement is presented in Figure 2.12. Note the initial increase in load to a local
maximum at which time the load drops to a lower level. Subsequently, the load varies
about a plateau load level throughout the crush process. The plateau region of the load-
displacement curve corresponds to progressive crushing of the crush tube. For the ﬁve

glass ﬁber crush tubes tested, the average SEA was 20.62 J/ gm. All testing of glass ﬁber
crush tubes reported in this chapter were conducted with the standard crush trigger. Plots

of the ﬁve glass ﬁber composite static crush tests are presented in Appendix A.

40

Table 2.5 Glass ﬁber composite crush tube SEA, quasi-static loading rate (0.0508

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

m/min).

Won Ave. Cross Max. Ave. Crush Total Energy SEA
Section Load (N) Load (N) Distance Absorbed (J) (J/ gm)

Area (cmz) (cm)
1A 9.95 41,350 32,140 12.7 4,082 17.96
”TEX 9.90 51,530 36,450 12.7 4,630 20.48
#32; 9.74 46,920 38,640 12.7 4,907 22.05
BB 9.90 45,830 38,260 12.7 4,860 21.49
133 9.91 49,940 37,680 12.7 4,785 21.14
Y—Ave. 9.88 47,1 10 36,630 12.7 4,653 20.62

 

Note: SEA based on material density of 1.8002 gm/cm3.

41

 

60.000 —

50,000

40,000 w

'0 30,000

20,000

1 0,000

Tube 3 #133

 

A

 

A11 AAA
vvvrvvvvv

WV’W

 

 

 

 

 

 

 

 

 

 

 

60 80

Displacement (mm)

42

120

Figure 2.12 Typical glass ﬁber composite quasi-static crush tube.

2.4.3 Carbon Fiber Composite Crush Tube Test Results
2.4.3.1 Carbon Fiber Composite Crush Tube Quasi-Static Test Results

The carbon ﬁber/Vinyl ester composite crush tubes were similar in geometry to
the glass ﬁber composite tubes but with a thinner wall thickness. The carbon ﬁber
composite tube outer section dimensions, presented in Table 2.2, were the same as the
glass ﬁber composite tubes but with a different overall length. Quasi-static test results
for the carbon ﬁber crush tubes are presented in Table 2.6. A typical quasi-static load-
displacement plot is presented in Figure 2.13. Plots of the ﬁve carbon ﬁber composite
quasi-static crush tests are presented in Appendix B. The appearance of the carbon ﬁber
composite tube quasi-static load-displacement curve is similar to that observed for the
quasi-static loaded glass ﬁber composite tubes. A plateau load follows an initial increase
in load with variation above and below the average plateau load.
2.4.3.2 Carbon Fiber Composite Crush Tube Dynamic Test Results

The dynamic test results for the carbon ﬁber composite are presented in Table 2.7.
A typical dynamic load-displacement plot is presented in Figure 2.14. Plots of the ﬁve
carbon ﬁber composite dynamic crush tests are presented in Appendix B. Note that the
average plateau load is lower for the dynamic loaded carbon ﬁber composite, 15,650 N,
than the same material loaded quasi-statically, 21,510 N. The average SEA for dynamic
loading was 23.8 J/gm vs. 30.4 J/gm for quasi-static loading. This is consistent with the
observation reported in literature search of Chapter 1 that dynamic loading results in a

lower SEA than static loading.

43

Table 2.6 Carbon ﬁber composite crush tube SEA, quasi-static loading rate (0.0508

 

 

 

 

 

 

 

 

m/min).
Coupon Ave. Cross Max. Ave. Crush Total Energy SEA
Section Load Load (N) Distance Absorbed (J) (J/gm)
Area (cmz) (N) (cm)

4278-2-1 5.26 22,300 18,170 12.4 2,250 24.25
4278-2-4 4.88 28,130 21,450 12.4 2,660 30.85
427 8-2-5 4.91 29,390 23,320 12.4 2,890 33.32
4278-2-9 4.93 27,590 22,260 12.4 2,760 31.66
427 8-2-17 4.92 29,210 22,330 12.4 2,770 31.87
Ave. 4.98 27,320 21,510 12.4 2,670 30.39

 

 

 

 

 

 

 

Note: SEA based on material density of 1.42 gm/cm3.

 

Tube 4278-2-17
Standard Trigger

 

30,000

..,... /\ /1 A

 

 

 

 

-.... I f” w MR7 V“ V W/

 

3......» [.7

 

 

 

I 0,000

 

 

 

 

 

 

 

 

 

 

 

0 20 40 60 80 100 120
Displacement (mm)

Figure 2.13 Typical carbon ﬁber tube quasi-static load-displacement plot.

45

Table 2.7 Carbon ﬁber composite crush tube SEA dynamic loading.

 

 

 

 

 

 

 

 

 

 

 

 

 

 

Coupon Load Ave. Cross Max. Ave. Crush Total SEA
Rate Section Load Load Distance Energy (J / gm)
(tn/sec) Area (cmz) (N) (N) (cm) Absorbed
(J)
4278-2-2 2.0 4.68 20,620 16,1 10 26.6 4,280 24.2
4278-2-3 2.0 4.84 22,060 16,100 26.8 4,330 23.5
4278-2-1 1 2.0 4.51 20,040 15,340 26.9 4,130 24.0
4278-2-12 2.0 4.53 20,890 15,030 27.3 4,100 23.4
427 8-2-16 2.0 4.65 22,130 15,680 27.1 4,250 23.8
Ave. 4.64 21,150 15,650 26.9 4,220 23.8

 

 

Notes: SEA based on material density of 1.42 gm/cm3 .

46

 

Load(N)

25.000

20,000

I 5.000

'09” "

5.000

Tube 4278-2-12

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

50 100 150 200 250
Displacement (mm)

Figure 2.14 Typical carbon ﬁber crush tube dynamic load-displacement plot.

47

300

2.5 Macroscopic Damage Modes

AS a tube is crushed, several modes of energy absorption are evident.
Delamination, ﬁber and resin fracture, damage due to bending, and sliding friction of the
composite against the plug trigger or platen were observed. Each of these energy
absorption modes absorbs varying amounts of energy depending on crush speed,
environmental conditions, and composite composition.

During a quasi-static or dynamically loaded tube crush test interlaminar shear
loads occur when the composite is forced to bend over the ﬁllet transition between the
plug and base plane of the plug trigger that is perpendicular with the crush tube axis and
crush direction. Delanrination occurs when layers, or plies, of composite separate due to
interlaminar shearing forces. Delamination can be seen in amulti-ply quasi-statically
loaded crushed glass ﬁber crush tube shown in Figure 2.15. Careful examination reveals
that there are six separate composite plies that have separated.

A photo of a quasi-statically compressed carbon ﬁber composite crush tube is
presented in Figure 2.16. A photo of a dynamically compressed carbon ﬁber composite
crush tube is presented in Figure 2.17. Careﬁrl examination shows two plies present in
the damaged carbon ﬁber composite consistent with the ﬁber architecture of the carbon
ﬁber composite. The appearance of the quasi-statically compressed and dynamically
compressed carbon ﬁber composite crush tubes are visually similar and suggest similar
energy absorption modes. These are: ﬁber fracture at the tube corners, delamination, and
matrix damage.

During crush, the damage process zone advances intermittently as ﬁbers are

broken and ﬁber bundles ﬁacture. Cracks advance to the next ﬁber or ﬁber bundle and

48

crack propagation is arrested until the building forces due to crush loading become
greater than the strength of the ﬁbers. Tensile forces pull the composite material apart at
the corners as the tube is forced over the trigger. The increasing perimeter of the trigger
that the composite tube is forced over loads the composite in tension similar to the tensile
loading imposed during a tensile test of a ﬂat coupon. In the case of the composite tube,
this tensile loading is perpendicular, or 90°, to the axis of the tube. The tensile strength
in the 90° material direction as reported in Table 2.1 for the glass ﬁber composite and
Table 2.2 for the carbon ﬁber composite is in the same material direction perpendicular to
the crush tube axis for both composites.

Damage to the composite due to bending was due to general matrix ﬁacture of the
composite as it was forced over the ﬁllets of the crush trigger. Fracture of the composite
occurs when the in-plane tension loads exceed the tensile strength of the composite. The
tensile strength of the composite is usually determined by the ﬁber strength, which is
signiﬁcantly higher than the resin matrix. As the composite tube sides are forced over the
trigger ﬁllets, bending causes tensile force in the 0° direction on the inside of the tube
adjacent to the trigger and compressive loading on the outside of the tube. Because the
ﬁber is signiﬁcantly stronger in tension than the matrix, the matrix cracks and
delamination occurs. The damage to the matrix prevents the transfer of load between
fibers and the ﬁbers are unloaded. This was observed for both glass ﬁber composite and
carbon ﬁber composite. As a result, there was ﬁber fracture at the tube comers where
high tensile loads were imposed. The bending loads were sufﬁcient to cause widespread
matrix fracture between ﬁbers with little ﬁber damage Observed in the tube sides

resulting in the loss of structural integrity of composite material.

49

Energy absorption due to ﬁiction was inferred by the physics Of the crush tube
test. It is readily observed that the composite is forced to slide over the trigger ﬁllets,
sides and base. Measurement of this energy absorption mode is the focus of this
research.

2.6 Microscopic Damage Modes

Matrix damage is mostly the result of bending of the composite as the composite
crush tube is forced against a ﬂat platen or over a plug trigger. As the composite is
crushed, bending forces cause the polymer matrix to fracture. The cracks found in the
matrix are closely spaced, typically the spacing between ﬁber tows. This is demonstrated
by the peaks of the load in the load-displacement plots presented in Appendix B. The
ﬁber tow spacing for both the glass ﬁber and carbon ﬁber composites was measured to be
4-5 mm. Examination of the load-displacement plots shows major load peaks that are
approximately 4 mm to 5 mm apart. Minor load peaks are closer together and are
probably due to individual ﬁber tow fracture at the tube comers. Major peak loads are
probably due to crack arrest at all four tube corners followed by near simultaneous ﬁber
tow fracture with an abrupt load drop at all four comers allowing the crack to propagate
to the next set of corner ﬁber tows where it arrests until the load builds to the ultimate
strength of all four fiber tows with the process repeating progressively as tube crush
continues.

Additional microscopic damage modes such as ﬁber/matrix debonding and
ﬁber/matrix pullout are probably present. These damage modes are not examined in this
research, as each microscopic damage mode may be present with or are a part of more

than one macroscopic energy absorption mode confounding efforts to isolate and measure

50

energy absorption by each microscopic damage mode. Such investigation is left to future

research.

51

 

Figure 2.15 Quasi-statically compressed glass ﬁber tube showing delamination, comer
splitting, and matrix damage due to bending.

52

 

Figure 2.16 Quasi-statically compressed carbon ﬁber tube showing delamination, comer
splitting and matrix damage due to bending.

53

 

Figure 2.17 Dynamically compressed carbon ﬁber composite crush tube showing
delamination, comer splitting and matrix damage due to bending.

54

CHAPTER 3. Characterization of Individual Damage Modes

3.1 Introductory remarks

It is necessary to identify the energy absorption modes and quantify their
magnitudes for a given crush rate before the energy absorption of a particular composite
material is understood. The ﬁrst part of this problem was addressed in Chapter 2 with
the careful observation of tubes subjected to crush loading. Energy absorbing modes
were identiﬁed using observation. The second part of the problem, quantiﬁcation of
energy absorption by mode, is more difﬁcult because of the coupling among energy
absorbing modes.

During crush, the peak load is determined by the force required to ﬁ'acture the
glass or carbon tows in the corner regions of the crush tube. At load levels below the
strength of the glass or carbon corner tows, other damage occurs in the composite tube
such as delamination and matrix damage due to bending. Energy absorption due to
sliding friction is dependent on the load normal to the sliding surface.

The local load maximums that occur during tube crush, such as shown in Figure
2. l 3, suggest that not all energy absorption modes are occuning simultaneously. If this is
so, it should be possible to isolate and measure energy absorption modes. Also, the
composite systems studied show linear or near-linear behavior.

It is the purpose of this research to determine if the relative energy absorption of
each mode can be separated and measured for both quasi-static and dynamic load rates.
AS diSeussed in Chapter 1, there is a strain rate effect that causes an increase in elastic
“109111118 and tensile strength of many composites. The effect described in Chapter 1 is

that for composite crush tubes the dynamic crush load and energy absorption are

55

signiﬁcantly reduced relative to static loading and remains relatively constant over a
range of dynamic load rates.

After energy absorption modes are identiﬁed, a test must be designed that will
isolate and measure each mode or groups of energy absorption modes. It is hypothesized
that the energy absorption due to ﬁiction mechanisms is one of the primary modes of
energy absorption for composite materials. It is further hypothesized that the difference
in ﬁiction energy absorption under quasi-static loading vs. dynamic loading is the
primary reason that the SEA of a composite varies depending on loading speed. Friction
will be addressed in detail in Chapter 5.

3 .2 Crush Tube Testing

3 .2.1 General SEA Calculation

Testing of crush tubes is accomplished by crushing the tube between the platens
of a test machine in a controlled manner. Displacement and load data obtained in this
type of test is used to calculate speciﬁc energy absorption (SEA). SEA is calculated by
integrating the load-displacement curve to determine total energy absorbed and then
dividing the total energy absorbed by the mass of the crushed portion of the tube. The
mass of the tube that is crushed is determined by calculating the volume of the tube that
is crushed and then multiplying the crush volume by the composite density. The crush
V01 ume was determined by multiplying the cross section area by the measured
displacement obtained during testing. The cross section area of the tube is calculated

usrng measurements of the cross section geometry.

56

3.2.2 Crush Triggers

Experience has shown that it is advantageous to use a crush trigger to start the
cruSh process of the composite. Ifno trigger is used, high loads can occur before
crushing begins resulting in little energy absorption due to structural instability or
overload. The maximum load for the glass ﬁber tube base on material compressive
strength was calculated to be 50.6 ken and 17.8 ken for the carbon ﬁber composite tube.
See Tables 2.2, 2.3, and 2.4 for tube characteristics and compressive strength for glass
and carbon ﬁber composite material. One type of simple crush trigger is a bevel at the
leading edge of the crush tube as Shown in Figure 3.1. For this study, a 45° bevel was
machined in one end of each crush tube shown in Figure 3.2. The crush process begins at
the point of the bevel at a relatively low load because of the small cross section. Once

the crush process zone is established the tube crushes progressively through the duration

Of the test at a relatively constant load.

57

 

Figure 3.1 Glass ﬁber composite tube showing bevel trigger.

58

A plug type trigger is frequently used with the beveled tube that results in the
crush process zone being established at a load plateau level below the tube buckle load.
A crush tube with a plug trigger in place is Shown in Figure 3.3. The crush trigger
consists Of a short steel plug that closely ﬁts inside the tube and a base perpendicular to
the axis of the tube with a ﬁllet at the intersection of the plug and crush trigger base. The
clearance between the composite crush tubes, both glass ﬁber and carbon ﬁber, was 1

mm. A drawing of the standard plug trigger is presented in Figure 3.4.

59

 

 

 

 

Figure 3.2 Crush tube bevel detail.

60

 

Figure 3.3 Glass ﬁber composite tube with bevel and plug trigger.

61

19.0 DIAJ-IOLE

 

 

 

 

>+——
><——

 

 

ALL DIMENSIONS IN MILUMETERS

 

 

 

 

 

 

 

 

L 44.650155
..... l‘ “W
F...\// 1*”
WM I'm

Figure 3.4 Drawing of standard plug trigger.

62

3.2.3 Crush Tube Energy Absorbing Modes

A typical glass ﬁber composite tube crush test specimen is presented in Figure
3.5 . A carbon ﬁber composite crush tube specimen is presented in Figure 3.6. Note that
there is ﬁ'acture of the composite at the comers and damage of the resin matrix. The ﬂat
portions of the tube between the corners have extensively damaged resin matrix resulting
in a loss of bending stiffness of the composite. Also Visible is delamination of the six

plies of the glass ﬁber composite tube in Figure 3.5 and two plies of the carbon ﬁber

composite tube Visible in Figure 3.6.

63

 

Figure 3.5 Typical quasi-statically compressed glass ﬁber composite tube.

.mw gong-era "“‘""’ ""'"—'“"~
. ﬂ . -‘ I.
’ :‘éﬁ HEEL? ‘0‘? 4 ‘44 a . .

Z;
“111.!

       

Figure 3.6 Typical quasi-statically compressed carbon ﬁber composite tube.

65

3.2.3.1 Corner splitting
The most apparent energy mode Observed is fracture of the crush tube at the four
tube comers. Examination of both the carbon ﬁber and glass ﬁber composite tubes
reveals fracture of the ﬁber tows occurs only at the four comers of the tube for both
quasi-static and dynamic loading. While the composite matrix is damaged extensively in
the crushed region of the tube, no fractured ﬁber tows were Observed in the ﬂat sides of
the tube. The damage to the composite at the comers was similar in appearance to the
damage observed when the composite was subjected to a standard tensile test.
Tensile testing was performed on both the glass composite and carbon composite
and reported in Chapter 2, Tables 2.3 and 2.4. The ﬁactures in the crush tube comers are

similar to those observed in the 90° tensile tests. The energy density at ﬁacture measured

at 0.791 J/cm3 for the glass ﬁber composite (Table 2.3) and 0.475 J/cm3 for the carbon

ﬁber composite (Table 2.4). Strain energy density is determined by mathematically
integrating strain over the affected material volume. Experimentally, the strain energy
density is determined by mathematically integrating strain energy over the affected
material volume.

The total corner splitting energy absorption of the crush tube is calculated by
multiplying the tube comer volume by the tensile strain energy density at fracture. The
SEA attributable to comer splitting energy is calculated by dividing total corner Splitting
energy by the total volume of the crushed tube. The comer splitting SEA is presented in

Table 3.1 for glass ﬁber and Table 3.2 carbon ﬁber composite.

66

Table 3.1 Glass ﬁber composite corner splitting SEA.

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

Tube # Comer Area Crush Fracture Tube Crush SEA
(ml) Length (cm) Ener Mass (gm) (J/gm)
(J/cm )

1A 2.54 12.7 0.791 227 0.112
2A 2.50 12.7 0.791 226 0.111
3A 2.49 12.7 0.791 223 0.112
33 2.50 12.7 0.791 226 0.111
133 2.50 12.7 0.791 227 0.111
Average 2.51 12.7 0.791 226 0.111

Table 3.2 Carbon ﬁber composite comer splitting SEA.
Tube # Corner Area Crush Fracture Tube Crush SEA
(cmz) Length (cm) Ener Mass (gm) (J/gm)

__ (J/cm )

; 4278-2-1 1.01 12.4 0.475 92.0 0.065
4278-2-4 0.90 12.4 0.475 85.4 0.062
4278-2-5 0.86 12.4 0.475 86.0 0.059
4278-2-9 0.91 12.4 0.475 86.3 0.062

\ 4278-2-17 0.97 12.4 0.475 86.1 0.066

T Average 0.93 12.4 0.475 87.2 0.063

 

 

 

 

 

 

67

 

 

3.2.3.2 Matrix Fracture Due to Bending
Matrix damage due to resin fracture is the second energy absorption mode that

was observed. Figures 3.5 and 3.6 Show that the ﬂat sides of the crush tube that have
been forced to follow the radius of the plug trigger and were deformed into a curved petal
shape. The stiffness of the petals has been reduced to the point that the crushed petals
can be deformed easily whereas the undamaged sides of the crush tube remain rigid. As
the ﬁbers in the petals are not fractured, the only composite component observed to be
damaged is the resin. Close examination of the damaged composite shows that the resin
has fractured into very small granules that remain attached to the ﬁber tow bundles. At
the edges of the composite petals, individual ﬁber tow bundles are observed to be free
ﬁom the petals. The ﬁber tows can be easily separated ﬁ'om the damaged composite
showing extensive damage Of the resin matrix and loss Of continuity between the ﬁber
tows. This characteristic of the damaged composite was consistent between the quasi-
static and dynamic tested specimens for both the carbon ﬁber and glass ﬁber composites.
A description and results of a new test method that quantiﬁes the speciﬁc energy
absorption attributable to matrix fracture and friction is presented in Chapter 5.

3.2.3.3 Delamination

Delamination was evident when the glass ﬁber composite and carbon composite
tubes were crushed. Six separate layers were easily identiﬁed for the glass ﬁber
composite tubes as shown in Figure 3.2. The carbon ﬁber composite tubes delarninated

into two layers. The number of surfaces created can be determined using the relation:

S=2(N-1)

68

Where: S is the number of surfaces
2 represents the two surfaces created after delarrrination

N is the number Of plies or layers

For the glass ﬁber composite, 10 surfaces were created. Two surfaces were created in the
carbon ﬁber composite by delamination.

Calculation of the energy absorbed by this damage mechanism was determined

using the Mode 2 critical strain energy release rate, One. The test procedure: End Notch

Flexure Testing of Fiber Reinforced Composites, General Electric Speciﬁcation NO

4013367-059, Issue 2, was used to measure the static value of Gnc. No testing was

conducted to determine the dynamic value of Gnc. The values of Gnc of the glass ﬁber
composite are presented in Table 2.3 and in Table 2.4 for the carbon ﬁber composite.
Multiplying Guc by the surface area that is created during tube crush and dividing by the

mass of the crush tube that is crushed results in a value for the SEA for this mode.

For example, the glass ﬁber composite tube #1A:

1. Gnc of glass ﬁber composite is 0.123 J/cm2 (see Table 2.3)

2. Crush distance is 12.7 cm (see Table 2.5)

E”

Tube section perimeter is 19 cm

4. Delaminated surface area is 2,409 cm2

S"

Cross section area is 9.95 cm2 (see Table 2.5)

Calculated crush volume equals 126 cm3

9‘

69

7. The density of glass ﬁber composite is 1.8002 grn/cm3

8. The calculated crush mass is 227.5 gm

9. The resulting delamination SEA is 1.31 J/gm

The results of these calculations are summarized and presented in Table 3.3 for
the glass ﬁber composite and Table 3.4 for the carbon ﬁber composite. Only static
testing for end notch ﬂexure (ENF) was performed for both glass ﬁber composite and

carbon ﬁber composite.

70

Table 3.3 Glass ﬁber composite SEA due to delamination.

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

Coupon Gnc Delaminated Szurface Crushed Tube SEA
(J/cmz) Area (cm ) Mass (gm) (J/gm)

1A 0.123 2,409 227 1.30

2A 0.123 2,406 226 1.31

3A 0.123 2,406 223 1.33

3B 0.123 2,406 226 1.31

1 BB 0.123 2,406 227 1.30

Average: 1.31

Note: SEA based on material density of 1.8002 gm/cm3.
Table 3.6 Carbon ﬁber composite SEA due to delamination.

Coupon G"c Delaminated Szurface Crushed Tube SEA
(J/cmz) Area (cm ) Mass (gm) (J/gm)

4278-2-1 0.164 451 92.5 0.798
4278-2-4 0.164 ' 450 85.9 0.857
4278-2-5 0.164 449 86.5 0.850
4278-2—9 0.164 450 86.8 0.848
4278-2—1 7 0.164 451 86.6 0.852
Average: 0.841

 

 

 

 

 

Note: SEA based on material density of 1.42 gm/cm3.

71

 

 

3.3 Tapered Trigger
3 . 3.1 Introductory Comments

The main difﬁculty in and quantifying energy absorption modes is developing a
simple test to separate energy absorbing modes. One of these approaches, the tapered
trigger, is discussed in this chapter. A new strip test is introduced and discussed in
Chapter 5 that was used to quantify energy absorption attributable to sliding friction and
energy absorption attributable to matrix damage due to bending.

When a tube is compressed quasi-statically or dynamically using the standard
plug trigger, the tube comers split, the ﬂat sides of the tube are forced follow the radius
of the trigger resulting in fracture of the resin matrix and delamination between layers of
the composite. The order that these damage modes occur was not determined. However,
once the crush process is started, all damage modes occur throughout the crush event. In
addition, there is energy absorption due to the friction of the composite sliding over the
trigger. The ﬁrst effort to isolate energy absorption modes was the development of the
tapered trigger. It was reasoned that if the composite tube is compressed using a small
angle trigger, delamination and matrix damage due to bending would be eliminated
leaving only corner splitting and sliding ﬁiction.

3 .3 .2 Tapered Trigger Design
The design features of the tapered trigger are presented in Figure 3.7. Note that
the tapered portion of the trigger is inclined 5 degrees relative to the axis of the trigger.
All tapered trigger tests were conducted using quasi-static load rate. During the crush
test, the composite crush tube is forced over the increasing perimeter of the tapered sides

resulting is tensile loads that cause the tube to fracture in the corners, similar to the

72

standard plug trigger. However, because the sides of the tube are not forced to bend
following the small radius between the standard plug trigger and its base, the composite
sides do not delaminate and the matrix resin does not fracture. This test effectively has
only two energy absorption modes, comer splitting and sliding ﬁiction. A tapered trigger
is presented in Figure 3.7 and a glass composite tube that has been tested using the
tapered trigger is shown in Figure 3.8. Note that the only visible damage is the fracturing

of the comers.

73

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

m N j
. w 5 am ..
w H mm as.
a... m mm m m
«a 5 mm mm s
4 r K
28:63:. mm f k
9.563.. mm 285093
36.8
33 Iv
V 28503» I 28:63.: I
6.8 4 goes #3 s , t
A Ali J
I I t IIIIIIIIIII
W. IIIIIIIIII
_
.
_
Fllllil|.'¢"'l""'

Figure 3.7 Tapered trigger drawing.

 

 

 

 

 

 

74

 

.7 Tapered trigger.

3

Figure

75

T. ‘1 .l'r F

In” ”I”. .15.

....... I

c 30.0 1..o..
Ha. o. a)
o r.

can...

..‘3.uu § . hind

a». ..

.4. ..r

 

Figure 3.8 Compressed glass ﬁber composite tube using tapered trigger.

76

3.3.3 Tapered Trigger Crush Tube Results
3.3.3.1 Glass Fiber Tapered Trigger Crush Tubes

The results Of tapered trigger crush tests on glass ﬁber composite tubes are
presented in Table 3.7. A load-displacement plot Of both of the tapered trigger specimens
is presented in Figure 3.9. Tube comer splitting begins at approximately 8 mm of test
machine crosshead travel after initial test free play is taken up. Note that tube TT-3
reached the ﬁrst load peak after 20 mm of crosshead displacement while tube TT-4
reached its ﬁrst load peak after only 10 mm of crosshead displacement. However, there
is a low load level observed for tube TT-3 until 12 mm of crosshead displacement while
only 2-3 mm of crosshead displacement corresponding to low level loading occurs in TT-
4 before loading increases. Recall that there is a 45° bevel at the leading edge of each
tube. The load level required to cause damage to the sharp leading edge Of the bevel is
low compared with the load required to cause damage to the full thickness in remainder
of the crush tube. This reduced load level at the beginning of the crush test is also due to
the smaller number of 90° ﬁbers of the glass ﬁber composite and 145° ﬁbers of the
carbon ﬁber composite that are continuous in the comer regions of the beveled end of the
tube. The 90° ﬁbers of the glass ﬁber composite and i45° ﬁbers of the carbon ﬁber
composite are the primary loadobearing component and in the beveled region of the tube,
these ﬁbers are mostly severed due to machining.

Individual plots of each test are presented in Appendix C. Examination of the
tested tubes showed that damage was conﬁned to the comers with no delamination or
damage of the resin in the ﬂat sides of the tubes. Average SEA using the tapered trigger

was 4.91 J/gm compared with 20.62 J/gm for a tube crushed using the standard plug

77

trigger. The large difference is because energy absorption attributable to matrix damage
due to bending is not present. When only two of the energy absorbing modes are present,

reduced load levels and energy absorption were observed.

78

Load (N)

14.000 -

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

M [JV
10.000 ,4 e M .

,IY '- .~

I A
8.000 .V' '0 '\"¢/\I‘_TZJ
"

6.000 -
4.000

-—-Tube TT3

- - - Tube TT4
2.000

0 .
0 20 40 60 80 100 120 140

Displacement (mm)

Figure 3.9 Glass ﬁber composite tubes, load-displacement plot for quasi-statically
loaded with tapered trigger.

79

Table 3.7 Glass ﬁber composite crush tube SEA using tapered trigger quasi-static load

 

 

 

 

 

rate (0.0508 m/min).
Coupon Ave. Cross Max. Ave. Load Crush Total Energy SEA
Section Load (N) (N) Distance Absorbed (J / gm)
Area (cmz) (cm) (J)
TT2 10.1 14,780 10,690 12.5 1,330 5.87
TT4 10.1 11,670 7,190 12.4 895 3.95
Ave. 10.1 13,225 8,940 12.5 1,113 4.91

 

 

 

 

 

 

 

Note: SEA based on material density of 1.8002 gm/cm3.

80

 

3.3.4.1 Carbon Fiber Tapered Trigger Crush Tubes

Results for the tapered trigger carbon ﬁber composite tube crush are presented in
Table 3.8. Two of the carbon ﬁber composite tubes that were crushed using the tapered
trigger are presented in Figure 3.10. Individual plots for each test are presented in
Appendix C. Examination of the tested tubes showed that damage was conﬁned to the
corners with no delamination or damage of the resin in the ﬂat sides of the tubes.
Average SEA using the tapered trigger was 7.442 J/gm compared with 30.39 J/gm for a
tube crushed using the standard plug trigger. As with the glass ﬁber composite tubes,
when only ﬁiction and comer splitting, reduced load levels and energy absorption are

observed.

81

Table 3.8 Carbon ﬁber composite crush tube SEA using tapered trigger quasi-static

 

 

 

 

 

 

 

loading rate (0.0508 m/min).
Coupon Ave. Cross Max. Ave. Crush Total SEA
Section Load Load (N) Distance Energy (J/ gm)
Area (cmz) (N) (cm) Absorbed
(J)

4278-2-13 4.863 9,564 5,729 12.4 713 8.296
4278-2-14 4.621 7,015 3,989 12.4 497 6.079
4278-2-15 4.866 8,878 5,345 12.4 665 7.735
4278-2-20 4.618 9,053 5,022 12.4 625 7.658
Ave. 4.742 8,628 5,021 12.4 625 7.442

 

 

 

 

 

 

 

Note: SEA based on material density of 1.42 grn/cm3 .

82

 

 

Figure 3.10 Crushed carbon ﬁber composite tube using tapered trigger.

83

3.4 Tapered Trigger Summary

The use of a tapered trigger to isolate comer splitting energy absorption provided
insight into the relative amounts of SEA attributable to the energy absorption modes
identiﬁed for square tube crush at quasi-static load rate. An energy balance for both the
glass ﬁber composite and carbon ﬁber composite crush tubes is presented in Table 3.9.
Table 3.9 shows that if energy superposition holds, splitting the comers of the tube
accounts for less than 1% of the total energy absorbed for both the glass and carbon ﬁber
composites.

This table also shows that delamination accounts for 6.4% of the energy absorbed
by the glass ﬁber composite and less than 1% of the energy absorbed by the carbon ﬁber
composite. This difference can be partially explained by the difference in the ﬁber
architectures. The glass ﬁber consists of six layers whereas the carbon ﬁber composite
contains only two layers. It was expected that the relative energy absorption due to
delamination of the carbon ﬁber composite would be less than that of the glass ﬁber
composite due to fewer layers present.

When the standard trigger is used, the balance of the SEA for the glass ﬁber
composite, 92%, and for the carbon ﬁber composite, 97%, is attributable to energy
absorption due to sliding friction and matrix damage due to bending. These ﬁgures were
alﬁved at based on the analysis given earlier and not by direct measurement. A unique
method of measuring energy absorption attributable to matrix damage from bending and
sliding ﬁiction is presented in Chapter 5. In addition, dynamic measurement of sliding

ﬁiction and matrix damage due to bending was also directly measured and is presented in

Chapter 5.

84

Table 3.9 Energy absorption summary for quasi—static loading.

 

 

 

 

 

 

 

 

Glass Fiber Composite Carbon Fiber Composite
Energy Absorption SEA SEA SEA SEA
Mode (J/gm) (%) (J/gmL 1%)
Corner splitting 0.111 0.54 0.063 0.21
Delamination 1.31 6.36 0.841 2.8
Friction 4.80 23.3 7.38 24.3
Remainder 14.38 69.8 22.11 72.7
(Matrix Damage)
Tube Crush 20.6 100 30.4 100
(Total)

 

 

 

 

 

85

 

CHAPTER 4. Analytical Solution to the Tapered Trigger Crush Test

4.1 Introduction

The tapered trigger test described in Chapter 3 was developed to isolate energy
absorption modes for a crush tube. It was demonstrated that corner splitting and sliding
friction were the most signiﬁcant energy absorption modes present.

The purpose in developing an analytical solution for the tapered trigger crush test is
to evaluate the reasonableness of the experimental results. It is also valuable to develop
insight into the physics of the crush process by modeling the crush tube and analytically
predicting its mechanical behavior using simple mechanics of materials concepts.

4.2 Free body diagram

A simple free body diagram of the tapered trigger tube crush test is presented in
Figure 4.1. A cross section of the composite tube at the beginning of the tapered plug is
presented with the applied load, Q, the normal force, N, and the tangential or ﬁiction
force, T. In addition, there is transverse shear force present because the composite is put
into bending as it is forced over the tapered trigger.

Note that the normal force is the resolved force from the distributed normal force,
n, over the length of composite ﬁ‘om the point where the crush tube is forced to begin
bending by the tapered trigger to the point at which the composite fractures as it is forced
by Q over the increasing cross section of the tapered trigger. The distributed normal
force is shown as having a triangular shape. This is consistent with the constant slope of
the tapered trigger. A distributed force can be represented as a point load passing through

the centroid of the distributed load. As the distributed load is a triangular shape, the

86

centroid of the distributed force is located 1/3 the distance from the base of the triangular

load distribution to the beginning of the distributed load.

87

 

Figure 4.1 Tapered trigger tube crush free body diagram.

88

The forces in the x-direction, parallel with the direction of the applied load, are
summed as follows to obtain an expression for the tangential force, T, and the normal

force, N.
ZFx = O

or,
0=Q-TcosO-Nsin0 ,
thus:
T=Q/cosG—Nsin0/cos0,
leading to :
T=Qsec0—Ntan0 eqn.(4.l)
The tangential force, T, is related to the normal force, N, by the relation:
T=pN
p is the dynamic coefﬁcient of friction between the ﬁberglass composite and the
steel trigger. This value for u was measured to be 0.24 using modiﬁed ASTM method
D1894-Ol "Standard Test Method for Static and Kinetic Coefﬁcients of Friction of
Plastic Film and Sheeting" Ref. (24), thus:
uN=Qse00—Ntan0
and:
N = Q sec 0/(u + tan 0) eqn. (4.2)

The forces in the y- direction are developed as:
2F y = 0

or,

89

=—V+NcosG—Tsin0
writing as:
V = N cos 0 - T sin 0 eqn. (4.3)

The moment, Mz , is determined at the point where the composite begins to
deﬂect due to the angle of the tapered trigger. The displacement distance, a, is the
distance ﬁ'om the point that the composite begins to deform to the point the maximum
value of the distributed force, n. The moment arm, 2a/3, is the distance from the point

that the composite begins to deﬂect to the centroid of the distributed force, n. Finally, the

moment is accounted for:
2M2 = 0

0 = M2 — V (2a/3)
or,
Mz = V (2a/3) eqn. 4.4

As the composite is forced over the increasing perimeter of the tapered trigger,
hoop forces are created perpendicular to the direction of the applied force and in the
plane of the composite. The maximum hoop force that is developed is limited by the
tensile strength of the composite in the hoop stress direction. This value for glass
composite used in this study presented in Table 2.3 as 189 MPa with corresponding
ultimate strain of 1.34%. At that point, the circumference of the composite tube has
increased to 101.34 % of its unstressed circumference. Refening to Figure 4.2, the width
of one side of the composite tube is L. The increase in length of one side of the tube at

the point of composite fracture is 25, the change in length of one side of the composite

9O

tube. The distance that the tube is pushed down the tapered trigger is a. The radius of the
comers, R, of the trigger is constant. For one side, the change in length is:
L + 25
and,

L+25=L+Zatan0

or,
8 = a tan 0
The total change in length is:
86 = 8 a tan 0
The total initial perimeter, P, is:
P = 4L + 2 1: R

R is the radius of the tapered trigger comers. Engineering strain is calculated by
dividing the change in perimeter length by the initial perimeter length:
8 = 8 8/P
a = 8 a tan 0/(4L + 2 n R) eqn. (4.5)
The crush distance, a, from the point at which the composite is ﬁrst subjected to
hoop forces sufﬁcient for corner splitting can be calculated by rearranging eqn. 4.5 as:

arm,x = emax(4L + 2 1r R)/8 tan 0 eqn. (4.6)

The measured, calculated and given values for the variables described above are

Presented in Table 4.1.

91

 

T

 

L+25

 

 

 

Figure 4.2 Section details for tapered trigger.

92

Table 4.1 Variable values.

 

 

 

 

 

 

 

 

 

 

 

 

 

 

Variable Name Variable Variable Value Value
Source
Plateau Load Q Measured 8,940 N
Plug Taper Angle 0 Measured 5°
Dynamic Coefﬁcient p, Measured 0.24
of Friction
Maximum Tangential Tm Calculated 6,576 N
or Friction Force
Maximum Normal NW1x Calculated 27,403 N
Force
Maximum Transverse Vm Calculated 27,872 N
Shear
Maximum Bending M Calculated 59.8 N m
Moment
Undeformed Tube Side L Measured 32.6 mm
Length
Maximum Change in 5m Calculated 0.282 mm
L/2
Tube Displacement at am“ Calculated 3.218 mm
Corner splitting
Tapered Trigger R Measured 6 mm
Comer Radius
Maximum Crush Tube 8m Measured 1.34%
Hoop Strain

 

 

 

 

93

 

The SEA attributable to friction can be calculated by dividing the product of the

tangential force, Tmax, (6576 N) and displacement with the mass of the tube crushed.
The cross section area of the glass ﬁber composite tubes is 10.11 cm2 and the crush

distance was 12.4 cm resulting in a total compressed volume of 125.4 cm3. The density

is 1.8002 gm/cm3 and the total crush mass was calculated to be 225.7 gm. The ﬁiction
SEA is calculated to be 3.65 J/gm.

In Chapter 3, the SEA attributable to sliding ﬁiction for the tapered trigger was
estimated to be 4.80 J/ gm or 98% of the total SEA. The analytical solution provides a
reasonable estimate of SEA due to sliding friction.

The SEA due to sliding ﬁiction with the standard plug was estimated in Chapter 3
to be 23% of the total SEA. Additional energy absorbing modes such as delamination
and matrix damage due to bending are observed during tube crush using the standard
trigger. An improved method of measuring sliding friction SEA for both static and

dynamic load rates is described in Chapter 5.

94

CHAPTER 5. Friction and Damage Process

5 .1 Lubricated Crush Tubes, Quasi-Static and Dynamic Loading

Based on the studies in the previous chapters, friction is believed to be a signiﬁcant
energy absorption mode for composite tube crush. A simple experiment was performed
to see if a change in the coefﬁcient of friction would result in a change in energy
absorption. The ﬁrst test was to apply a low friction coating in the interior surface of
two carbon composite crush tubes. These tubes were the same material, geometry, and
ﬁber architecture as those tested and reported in Chapter 2. The coating selected was
SprayonTM $00312 dry powder zinc stearate mold release. The coating is supplied in a
spray can and applied similarly to spray on paint. The coating dries immediately to a
white, dry, slippery ﬁlm. Two crush tubes were then loaded quasi-statically (0.508
m/min) in the same manner as the tubes discussed in Chapter 2. The results of these tests
are presented in Table 5.1 and with load-displacement plot presented in Figure 5.1. The
results show average SEA of 18.0 J/gm compared with 30.4 J/gm (Table 2.6) for carbon
ﬁber composite tubes loaded quasi—statically without zinc stearate coating, a reduction of
12.4 J/gm, over 40%. The damage appearance of the crushed zinc stearate treated tube
was similar to the untreated tube. This result suggests that energy absorption due to

sliding friction is signiﬁcant and can be isolated from SEA attributable to delamination

and matrix damage due to bending.

95

Table 5.1 Carbon composite crush tube SEA with zinc stearate coating at quasi-static

 

 

 

 

loading (0.0508 m/min).
Ff Coupon Ave. Cross Max. Ave. Load Total SEA
Section Load (N) Load (N) Distance Energy (J/gm)
Area (cmz) (cm) Absorbed
(J)
4278-2-7 5.26 15,870 11,420 1 1.4 1,302 17.4
4278-2-10 4.48 14,910 11,770 1 1.4 1,340 18.5
Ave. 4.87 15,390 11,595 11.4 1,321 18.0

 

 

 

 

 

 

 

 

 

Note: SEA based on material density of 1.42 grn.cm3

96

The appearance of the load-displacement curves was different between the two
specimens that were tested. The curve of tube #4278-2-7 displays local peaks with a
difference between the local minimum and maximum loads of approximately 4,000 N
that were approximately 20 mm apart. Compare with approximately 1,000 N between
the local maximum and minimum loads that were approximately 5mm between load
peaks displayed for tube #4278-2-10. One possible explanation for this difference is the
damage process between the two tubes was different in the way the :45 ﬁbers were
ﬁ’actured. If the ﬁactures advanced by fracture of the :45 ﬁbers in the corners

simultaneously and then arrested at the next set of :45 ﬁbers, the load would decrease to
a much lower level than in the case where the corner splitting advanced sequentially and
the load-displacement diagram would show a higher frequency. Evidence for this
explanation can be seen by noting that the difference between local maximum and
minimum loads was 4,000 N with approximately 20 mm between load peaks for one tube
and for the other tube 1,000 N between local maximum and minimum loads with
approximately 5 mm between load peaks. The appearance of the damaged specimens
was identical and the average loads were similar. The difference in damage behavior

between the two tubes is likely normal variation in the architecture of the ﬁber

reinforcement.

97

18.000 ~

16.000

 

 

Local Maxirnurns

 

 

l

\

 

o"'

A

;

A

 

14.000

 

1

 

v \

 

 

 

 

 

J—427a-2-7

 

 

 

 

 

 

 

 

 

 

1 ----- 4270-2-10

 

 

 

 

 

 

 

 

 

20 40

 

60

Displacement (mm)

 

 

100

120

Figure 5.1 Carbon ﬁber composite tube crush quasi-static (0.0508 m/min) coated with
zinc stearate.

98

Dynamic testing was conducted on carbon ﬁber composite crush tubes lubricated
with zinc stearate coating. The results of this testing are presented in Table 5.2 and
Figure 5.2. Note that the test conducted at 500 mm/sec had SEA similar in value to the
two tests conducted at 2.0 m/sec with similar load-displacement curves. This shows that
for these two load speeds, the SEA is insensitive to load rate. The dynamically loaded
SEA for this group was slightly higher than the quasi-statically crush tubes lubricated
with zinc stearate (21.8 J/gm vs. 18.0 J/gm). It is important to note that the quasi-
statically loaded tubes were not coated with zinc stearate at the same time as the dynamic
loaded tubes presented in Table 5.2. Zinc stearate lubricant was applied from a hand held
spray can and application quality was evaluated visually. It is possible that the
application of lubricant was not consistent between these two. groups of specimens and
the coefﬁcient of friction was different between the two sets of tests. It is important to
note that the application of zinc stearate lubricant did make a difference in energy
absorption in both non-lubricated, quasi-statically and dynamically loaded carbon
composite crush tubes. A summary of the zinc stearate lubricated tubes vs. the non-

lubricated tubes is presented in Table 5.3.

99

Table 5.2 Carbon ﬁber composite crush tube SEA with zinc stearate low ﬁiction coating,

 

 

 

 

 

dynamic loading.
Coupon Load Ave. Cross Max. Ave. Load Total SEA
Rate Section Load Load Distance Energy (J / gm
(In/sec) Area (cmz) (N) (N) (cm) Absorbed )
(J)

4278-2-6 2.0 4.66 20,450 15,090 26.9 4,052 22.0
4278-2-18 2.0 4.53 18,680 13,830 26.8 3,709 21.5
4278-2-19 0.5 4.84 20,660 15,100 27.0 4,082 22.0

Ave. -- 4.68 19,930 14,673 26.9 3,948 21.8

 

 

 

 

 

 

 

 

 

Notes: SEA based on material density of 1.42 gm/cm3.

Table 5.3 Carbon tube test summary non-lubricated vs. lubricated with zinc stearate.

 

 

 

 

 

 

r SEA (J/gm) SEA (J/gm)
Quasi-Static Loading Dynamic Loading
r Without Zinc Stearate 30.4 23.8
With Zinc Stearate 18.0 21.8

 

100

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

25,000.00
20,000.00 3 . ' '
i ‘, ' t
" 3: I r. 1‘ ' ii é I
, ,r :1 32p"?- :1. - "I". ‘ .
15,000.00 - .' - . . ° - ' ' '
i .‘ ‘ i '. . l ‘
i ‘
10,000.00 ll - '
Iv — -4278-2-6 2m/sec
——4278-2-18 2mlsec
5.000.00 ----- 4278-2-19 0.5 m/sec
0.00 .
o 50 100 15° 2°° 25° 30°

Displacement (mm)

Figure 5.2 Carbon ﬁber composite tube crush dynamic loading coated with zinc stearate.

101

5.2 Strip Testing

5.2.1 Strip Test Introduction

Crush tubes lubricated with zinc stearate has demonstrated that altering the
coefﬁcient of ﬁiction signiﬁcantly affects the SEA of quasi-statically loaded carbon ﬁber
composite crush tubes. What this experiment did not do was to quantify the SEA
attributable to ﬁiction. What is needed is a test that isolates ﬁiction by removing it from
the crush process. To accomplish this goal, a test ﬁxture was proposed to test strips of
composite material that would remove friction energy absorption from the crush process
present when carbon composite is forced to slide over the steel surface of a crush trigger
for quasi-static and dynamic load rates. The test ﬁxture was designed to simulate the
damage process present in the sides of the crush tube when it is compressed using the
standard plug trigger described previously.

Figure 5.3 presents a schematic of the strip test ﬁxture. The main working parts
are four rollers mounted in low friction ball bearings. Because rollers are supported by
ball bearings, there is no sliding ﬁiction present when the rollers rotate as the strip
specimen is forced downward. Three guide rollers that support the test strip are of the
same diameter, 12.7 mm. The design and fabrication was simpliﬁed by keeping the
diameters of the three top rollers the same. The diameter of the three rollers was selected
because the roller radius is similar to the 6 mm radius of the standard plug trigger used in
the tube crush test. Rollers A and C are aligned in the vertical direction. The diameter of
the top two rollers, rollers, A and B, was minimized to keep the test ﬁxture compact and
minimize the length of the strip specimen. The roller second from the top, roller B, is

adjustable horizontally to accommodate varying thicknesses of samples. The lowest

102

roller, roller D, is twice the diameter, 25.4 mm, of the other three rollers and is also
horizontally adjustable. The larger diameter of roller D was chosen so that the range of
curvature imposed on a strip test specimen could be varied over a large range by
adjusting the horizontal position to simulate the curvature imposed on crush tubes when
triggers with ﬁllets of varying radii are used. Roller D is positioned to force the
composite strip specimen to bend against the bottom guide roller, roller C, as the
specimen is forced downward during testing.

Roller D was designed to have two test modes, locked and unlocked. When
unlocked, roller D is free to rotate. When locked, roller D is prevented from rotating with
a pin that ﬁts into a hole in the roller that prevents the roller from rotating forcing the
strip specimen to slide over roller D. The difference in the manner of loading from the
unlocked or free mode is due only to the sliding ﬁiction present when the roller D is kept
from rotating, forcing the strip to slide over the surface of the roller D. Delamination and
matrix damage due to bending are present for both locked and free roller D conditions.

Sliding ﬁiction is removed from the process when the roller D is free.

103

E

/— [A]12.7DlA.ROU.ER-FIXED

20.0 /— [B] 12.7 DlA. ROLLER - ADJUSTABLE
h
’1 V'Ir 5‘1}.

V ALLA ,, -- +
A it ‘a . " -.. .r .r

 

 

 

  
  

 

 

 

 

209 /—— [c1 iZ.7DiA.ROLLER-FD(ED
V "
A
20.0 /—— [DIZSADIAROLLER-ADJUSTABLE
v t “I ‘ -“ .,..
ALL DIMENSIONS 1N MlLUMEl'ERS

Figure 5.3 Strip test ﬁxture schematic.

104

5.2.2 Strip Test Specimen and Fixture Description

Test specimens were machined from a ﬂat plaque made from the same material as
the crush tubes but with a thicker gage (4.0 mm vs. 2.0 m) that was fabricated by
doubling the number of layers of braided carbon tows. The ﬂat plaque was fabricated for
material characterization and needed greater thickness for some of the mechanical tests
such as short beam shear test, end notch ﬂexure test, and compression tests. Several strip
specimens were machined ﬁom crush tubes and were tested in the strip test ﬁxture. The
thinner strip could not be tested due to structural instability. The portion of the thin strip
above the rollers buckled as soon as the end of the strip contacted roller D and the load
began to increase.

The strip dimensions are nominally 25.4 mm wide and 200 mm long. One end of
each strip is beveled at a 45° angle similar to the bevel present in the crush tubes
described previously. The purpose of the bevel is the same as the bevel in the crush
tubes, to lower the initial load peak at the start of the test and improve structural stability
of test. When the strip is compressed, the bevel is oriented away from the roller D, or to
the left as shown in Figure 5.3. A strip test specimen is shown in Figure 5.4. The strip
test ﬁxture with a strip specimen ready to be tested is presented in Figure 5.5. A typical

crushed specimen is presented in 5.6.

105

 

- Strip Coupon
' Roller A

Roller B (hidden)
Roller C
Roller D

    

 

Figure 5.6 Typical tested strip coupon.

107

5.2.3 Quasi—Static Strip Tests
Five strip specimens were quasi-statically (0.0508 m/min) loaded with the roller D
locked. The results of these tests are presented in Table 5.4. Five additional strip
specimens were quasi-statically (0.0508 nr/min) loaded with the roller D locked and the
strips lubricated with zinc stearate. The results of these testes are presented in Table 5.5.
A third set of ﬁve strip specimens were quasi-statically (0.0508 m/min) loaded with the
roller D free to rotate during testing. The results of these testes are presented in Table
5.6. Similar to the crush tube test results, the load-displacement curve was numerically
integrated to calculate the absorbed energy. The calculated absorbed energy was divided
by the mass of the portion of the strip that was compressed to determine the SEA for each
specimen tested. A typical load-displacement plot is presented in Figure 5.6. Note that
there is a high peak load with load drop-off to a plateau load. No damage to the
composite strip occurs until a peak load is achieved at which time the damage process
begins in the strip. Once damaged, the load required to create new damage in the tube
occurs at a lower level. However, in the strip test there is no need for the load to become
large enough for ﬁber ﬁacture as in the crush tubes. All plots for the specimens listed in
Tables 5.4, 5.5, and 5.6 are presented in Appendix D. A discussion of these results is

presented in section 5.2.5.

108

Table 5.4 Carbon ﬁber composite strip SEA, locked roller D quasi-static loading (0.0508

 

 

 

 

 

 

 

 

m/min).
Coupon Cross Max. Ave. Load Total SEA
Section Load (N) Load (N) Distance Energy (J / gm)
Area (cmz) (cm) Absorbed
4374-4-3-V 1.19 7,035 3,414 11.5 3%)] 20.1
43 74-4—3-W 1.18 7,201 3,295 11.4 377 19.5
43 74-4-3-X 1.05 4,285 2,815 1 1.4 322 18.8
4374-4-3-Y 1.18 6,115 3,170 11.4 368 18.8
4374-4-3-Z 1.15 7,961 3,147 11.4 360 19.1
Ave. 1.15 6,519 3,168 11.4 364 19.3

 

 

 

 

 

 

 

Notes: SEA based on material density of 1.42 gm/cm3.

109

 

Table 5.5 Carbon ﬁber composite strip SEA, locked and lubricated roller D, quasi-static

 

 

 

 

 

 

 

 

loading (0.0508 nr/rrrin).
Coupon Cross Max. Ave. Load Total SEA
Section Load (N) Load (N) Distance Energy (J/gm)
Area (cm) Absorbed
(00121 (J)

4374—4-3-AA 1.16 7,344 2,534 11.4 290 15.3
4374-4-3-BB 1.19 6,685 2,553 11.5 292 15.1
43 74-4-3-CC 1.16 7,041 2,400 l 1.5 275 14.5
43 74-4-3-DD 1.14 6,973 2,420 1 1.5 277 14.9
4374-4-3-EE 1.16 6,719 2,273 11.5 260 13.7
Ave. 1.16 6,952 2,436 1 1.5 279 14.7

 

 

 

 

 

 

 

Notes: SEA based on material density of 1.42 gm/cm3.

Strip lubricated with zinc stearate dry ﬁlm.

110

 

Table 5.6 Carbon ﬁber composite strip SEA, free roller D,
quasi-static loading (0.0508 m/min).

 

 

 

 

 

 

 

 

Coupon Cross Max. Ave. Load Total SEA
Section Load (N) Load (N) Distance Energy (J/grn)
Areza (cm) Absorbed
(cm ) (J)

4374-4-3-FF 1.16 5,782 2,133 11.5 244 12.9
4374-4-3-GG 1.16 5,996 2,417 11.5 277 14.6
43744-3-HH 1.16 5,887 2,171 11.5 249 13.1
43 74-4-3-11 1.16 6,332 2,363 11.5 271 14.3
4374-4-3-JJ 1.13 2,556 1,881 11.5 215 11.7
Ave. 1.15 5,311 2,193 11.5 251 13.3

 

 

 

 

 

 

 

Notes: SEA based on material density of 1.42 gm/cm3.

lll

 

Strip 43744-3.v

 

 

 

 

 

g 4:000 V '\v/\\V VAX/1%
3
3 ’OOO - RWW

 

 

 

 

 

 

 

 

 

 

 

 

 

0 20 4O 60 80 100 120

Displacement (mm)

Figure 5.7 Typical load-displacement plot for quasi-static loaded
(0.0508 m/min) carbon composite strip with locked roller D.

112

5.2.4 Dynamic Strip Tests
Five strip specimens were tested dynamically at 2.0 m/sec with roller D locked,

and three additional specimens were tested at 0.5 m/sec with the roller D locked. The
results of these tests are presented in Tables 5.7 and 5.8. There was no signiﬁcant
difference in the SEA of strips tested at 2.0 m/sec and 0.5 m/sec (15.5 J/gm vs. 15.4

J/ gm). This shows that for these two load rates that the carbon composite is load rate
insensitive. A plot of a typical load-displacement curve for 2.0 m/sec and 0.5 m/sec
loading is presented in Figure 5.7. The curves are similar with no obvious difference in

appearance. The curves for all of the specimens described in Tables 5.7 and 5.8 are

presented in Appendix D.

113

Table 5.7 Carbon ﬁber composite strip SEA, locked roller D, dynamic loading (2.0

 

 

 

 

 

 

 

 

m/sec).
Cross Total
Coupon Section Max. Load Ave. Load Energy SEA
Load Drstance
Area (N) (N) (cm) Absorbed (J / gm)
(cu?) (J)
4374-4-3-QQ 1.07 6,607 2,253 10.9 246 16.6
4374-4-3-RR 1.12 8,194 2,429 10.8 263 15.7
4374-4-3-SS 1.1 1 10,307 2,322 1 1.0 256 14.6
4374—4-3-TT 1.13 6,882 2,243 10.9 244 13.8
4374-4-3-UU 1.04 5,035 2,513 10.9 275 16.9
Ave. 1.09 7,405 2,352 10.9 257 15.5

 

 

 

 

 

 

 

Notes: SEA based on material density of 1.42 gm/cm3.

Table 5.8 Carbon ﬁber composite strip SEA, locked roller D, dynamic loading (0.5

 

 

 

 

 

 

m/sec).
Coupon Cross Max. Ave. Load Total Energy SEA
Section Load Load Distance Absorbed (J) (J/ gm)
Area (cmz) (N) (N) (cm)

4374-4-3-BBB 1.08 6,897 2,139 1 1.1 237 13.9
4374-4-3-CCC 1.12 6,958 2,503 10.9 273 15.6
4374-4-3-DDD 1.08 7,706 2,579 1 1.0 285 16.7
Ave. 1.09 7,187 2,407 l 1.0 265 15.4

 

 

 

 

 

 

 

Notes: SEA based on material density of 1.42 gm/cm3 .

114

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

9,000
0.000 ~—.’——
1
‘1
7.000 1'
-"-4374-4-3-RR2mI8 :
1—4374-4-3-cccosmi
0,000 ...............................
g15,000 .
3 I
4.000 '
l ,‘
° ". n l ‘
3000<~ i ~° ' Lt . iii
' \ g, 1
1' | r.
1 -_1 | / .. .
1' , , . .
2.” 1 ‘ " 1 v ‘.1 ‘ ' \'.|.1'” "..‘“. .
k V. 7 ."r
1,ooo«~———- %
0 A.
0 20 40 00 80 100 120
DlepiacemenNmm)

Figure 5.8 Typical carbon composite strip load-displacement dynamic loading with
locked roller D.

115

Roller D was locked and specimens were loaded at 2.0 m/sec for the next set of
ﬁve strip specimens. However, these ﬁve strips were coated with zinc stearate to reduce
sliding friction. The results of these tests are presented in Table 5.9. The curves of the

tests are presented in Appendix D.

116

Table 5.9 Carbon ﬁber composite strip SEA, locked and lubricated roller D, dynamic

 

 

 

 

 

 

 

 

loading (2.0 nr/sec).
Coupon Cross Max. Ave. Load Total Energy SEA
Section Load Load Distance Absorbed (J) (J/gm)
Area (cmz) (N) (N) (cm)
4374-4-3-VV 1.13 6,592 2,132 10.9 233 13.2
43 74-4-3-WW 1.09 8,453 2,095 10.6 222 13.4
4374-4-3-XX 1.10 8,850 2,284 11.0 249 14.6
4374-4—3-YY 1.10 10,468 2,278 1 1.0 251 14.5
4374-4-3-AAA 1.19 8,835 2,305 10.8 249 14.6
Ave. 1.12 8,640 2,219 10.9 ' 241 14.1

 

 

 

 

 

 

 

Notes: SEA based on material density of 1.42 gm/cm3.

Lubricated with zinc stearate dry ﬁlm.

117

 

The results for dynamic loading of strips with roller D free to rotate are presented
in Table 5.10 for 2.0 m/sec and in Table 5.11 for 0.5 m/sec. A typical plot with load-
displacement curves for 2.0 m/sec loading and 0.5 nr/sec is presented in Figure 5.8. The
appearance of the curves is similar and the average SEA for 2.0 m/sec and 0.5 m/sec are
close in value (13.0 J/gm vs. 13.3 J/gm). Plots of all strip specimens shown in Tables

5.10 and 5.11 are presented in Appendix D.

118

Table 5.10 Carbon ﬁber composite strip SEA, ﬁ'ee roller D, dynamic loading (2.0

 

 

 

 

 

 

 

 

 

 

 

 

 

 

m/sec).
Coupon Cross Max. Load Ave. Crush Total SEA
Section (N) Load Distance Energy (J / gm)
Area (N) (cm) Absorbed
<ch (J)

4374-4-3-LL 1.13 6,546 1,934 10.9 210 11.9
4374-4-3-MM 1.13 9,659 2,073 10.9 227 12.8
4374—4-3-NN 1.05 7,706 2,081 10.9 228 13.8
4374-4-3-OO 1.09 7,385 2,170 10.9 237 13.9
4374-4-3-PP 1.10 6,119 2,018 10.9 220 12.8
Ave. 1.10 7,479 2,055 10.9 224 13.0

 

Notes: SEA based on material density of 1.42 gm/cm3.

Table 5.11 Carbon ﬁber composite strip SEA, free roller D, dynamic loading (0.5 m/sec).

 

 

 

 

 

 

 

 

 

 

 

 

Coupon Cross Max. Ave. Crush Total SEA
Section Load Load Distance Energy (J/gm)

Area (cmz) (N) (N) (cm) Absorbed

(J)

4374-4-3-EEE 1.05 2,960 1,794 11.0 198 12.0
4374-4-3-FFF 1.09 5,402 2,317 11.1 256 14.8
4374-4-3-GGG 1.09 5,493 2,038 11.1 226 13.1
Ave. 1.08 4,618 2,050 11.1 227 13.3

 

Notes: SEA based on material density of 1.42 gm/cm3.

119

 

 

Dynamic Crush 2 mlsec
Strip 43744-34411
Free Roller

 

O 20 40 60 80 100 120
Displacement (mm)

Figure 5.9 Typical strip load-displacement dynamic loading and free roller D.

120

5.2.5 Strip Test Results Discussion and Summary

A summary of strip testing is presented in Table 5.12. The variables presented
are: load speed, roller D ﬁxity, and the presence or absence of lubrication. The ﬁrst
comparison can be made for cases with varied load speed, locked roller D, and specimens
lubricated or not lubricated with zinc stearate. When the loading is quasi-static (case 1),
the average SEA is 19.3 J/m. When the loading is dynamic, either at 2.0 m/sec (case 4)
or 0.5 m/sec (case 7), the average SEA is 15.4 J/m, a 20% reduction compared with the
quasi-static result.

When the roller D is locked and the strips are lubricated with zinc stearate, the SEA
for quasi-static loading (19.3 J/grn of case 1 vs. 14.7 J/gm of case 2) vs. dynamic loading
at 2.0 m/sec (14.1 J/ gm of case 5). The zinc stearate lubrication resulted in a reduction in
SEA of 23.8% (14.7 J/gm of case 2 vs. 19.3 J/gm of case 1) compared with quasi-static
loading. However, when tested dynamically at 2.0 m/sec, the reduction in SEA for the
lubricated strips was only 8.4% compared with the non-lubricated test strips (14.1 J/gm
of case 5 vs. 15.4 J/ gm of case 4). This was the ﬁrst test that shows that sliding ﬁiction
absorbs smaller relative amount of energy when a composite tube is crushed dynamically.
The difference may also be attributed to the relatively reduced roll that ﬁiction plays in
dynamic loading cases than in the quasi-static cases. The difference in laboratory
facilities may also be a factor . The static strip testing was performed at the Ford
Scientiﬁc Research Laboratories whereas the dynamic testing was conducted at the
National Transportation Research Center at Oak Ridge National Laboratory, Tennessee.
The application of the zinc stearate dry ﬁlm lubricant was with a hand-held spray can and

the application may have differed between the two sets of strip specimens.

121

When the roller D is free to rotate there was little difference observed in the SEA
between quasi-static loading (13.3 J/gm of case 3) and dynamic load rate of 0.5 rn/sec
(13.3 J/gm of case 8). There was only a slight difference in SEA when the dynamic load
rate is 2.0 m/sec (13.0 J/gm of case 6). This result and the results described in the

previous two paragraphs clearly demonstrate that:

l. Sliding friction absorbs a signiﬁcant amount of energy at least in quasi-static
cases.

2. Altering the coefﬁcient of friction of the composite tube surface can vary
SEA.

3. Energy absorption due to sliding friction can be separated from those due to
damage formation.
4. Energy absorption due to delamination and bending was affected by loading

rate for the load rates tested.

122

Table 5.12 Strip test results summary.

 

 

 

 

 

 

 

 

 

 

 

 

Case Crush Speed Roller D Lubrication Average SEA
(J/gm)
1 Quasi-Static Locked None 19.3
2 Quasi-Static Locked Yes 14.7
3 Quasi-Static Free None 13.3
4 2.0 m/sec Locked None 15.4
5 2.0 m/sec Locked Yes 14.1
6 2.0 m/sec Free None 13.0
7 0.5 m/sec Locked None 15.4
8 0.5 m/sec Free None 13.3

 

 

 

 

 

123

 

5.3 Energy Balance

The energy absorption modes described in Chapter 3 were partially quantiﬁed using
simple mechanical tests. Speciﬁcally, comer splitting and delamination were determined
to be minor energy absorption modes. Comer splitting accounted for only 0.21% (0.063
J lgrn) of the quasi-static SEA. Delamination accounted for 2.7% (0.84 J/ gm). The
remaining 97%, 29.5 J/ gm (Table 3.9) was attributed to sliding friction and matrix
damage due to bending. A tapered crush plug was used to obtain an inferred energy
absorption value of 24.3 %, 7.38 J/gm (Table 3.9) for friction with the balance of 72.7%,
22.11 J/gm (Table 3.9) attributable to bending damage. There was no matrix damage or
delarrrination observed in any of the tubes tested using the tapered trigger. The strip test
results provide an improved estimate of energy absorption due to sliding friction and
bending.

The energy absorbed in the strip test with quasi-static loading and locked roller D
was 19.3 J/gm (Table 5.13). The SEA attributed to sliding ﬁiction and damage due to
bending in a carbon composite crush tube was 29.5 J/ gm (Table 3.9). The difference of
10.2 J/gm can be attributed to the reduced damage observed in the strips. The appearance
of the damaged strips does not show the degree of matrix damage seen in tubes. A
limitation of the strip test is the larger radius that the composite strip is forced over during
the strip test compared with the tube crush tube. The strip test ﬁxture is limited to
bending the strip over a minimum radius of 6.3 mm, the radius of roller C. In addition,
the test specimens were twice the thickness of the composite crush tubes (4 mm vs. 2
mm). The deformation radius of the outside of a strip is at least 10 mm when the

minimum inside radius is 6 mm. When the carbon composite tube is forced over the

124

standard trigger, the maximum bending radius of the composite is the radius of the plug
trigger, 6 mm. The minimum bending radius is smaller by the thickness of the 2 mm
thick composite, or 4 mm. Thinner specimens could not be tested with the strip test
ﬁxture due to test coupon structural instability. A 2 mm carbon strip coupon was tested
with poor results. The thin test specimen buckled above the test ﬁxture. The test ﬁxture
was designed with the capability of testing various width specimens. As a result, the
diameter of rollers A, B, and C could not be fabricated with the required diameter, 8 mm,
correct for 2 mm thick strips and yet be stiff and strong enough to test thicker strips. A
reﬁned design of the strip test ﬁxture that contains guides above roller A and with
reduced width could yield improved test data. The important information obtained from
the strip test described in this thesis is the ratio of SEA due to sliding ﬁiction compared
with SEA due to bending.

There is one additional adjustment to the calculation of friction and bending SEA
that must be made due to the increased thickness of the strip used in the strip test. The
material used for strip testing was taken ﬁ'om a plaque that was twice the thickness
containing twice the number of braided plies. This means that there are three
delamination planes compared with the one delamination plane of the carbon composite
tubes. Thus, the SEA value for delarrrination for the strip test is 2.52 J/gm or three times
the delanrination value for carbon composite tubes, 0.84 J/gm (Table 3.6). This
delamination SEA value of 2.52 J/ gm was subtracted from the quasi-static load rate,
locked roller D strip test SEA of 19.3 J/m (Table 5.12), giving a sliding friction plus
bending damage SEA of 16.8 J/grn. The sliding friction and bending SEA values were

then calculated and tabulated in the second column of Table 5.13.

125

The ﬁiction and bending SEA for the carbon composite tubes loaded quasi-statically
was determined to be 97% of the total SEA or 29.5 J/m. This value was divided by the
strip friction plus bending damage SEA of 16.8 J/gm to obtain a ratio to quantitatively
determine the ﬁiction and bending SEA for carbon composite tube crush. These values
are presented in the third column of Table 5.13. Because ﬁiction is dependent upon the
crush load which is affected by the bending strength as well as the delamination and
comer splitting forces required for progressive crush, the ratio of sliding friction SEA to
bending SEA is assumed to not vary between the strip test results and the tube crush
results. Note that friction is determined to account for 10.6 J/gm or 34.8 % of the energy
absorbed when the carbon composite tubes are crushed at the quasi-static rate.

Determination of the dynamic crush energy balance was determined in a similar
manner and is presented in Table 5.14. The energy absorption due to bending was
essentially unchanged (10.78 J/gm vs. 10.43 J/gm). Dynamic energy absorption due to
sliding friction decreased signiﬁcantly, 4.26 J/gm vs. 10.6 J/gm for quasi-static loading.
Compare this reduction in SEA due to sliding friction of 6.34 J/gm with the overall
reduction in carbon composite tube crush SEA from 30.4 J / gm for quasi-static load rate
to 23.8 J/gm for dynarrric load rate, a reduction of 6.6 J/gm. It can be concluded that for
this composite system, nearly all of the measured SEA reduction when the load rate is
changed from quasi-static to dynamic, 2.0 m/sec, is attributable to sliding ﬁiction. A

summary is presented in Table 5.15.

126

Table 5.13 Carbon composite strip and crush tube energy balance quasi-static loading.

 

 

 

 

 

 

 

 

 

 

 

 

 

 

Energy Mode Strip Tube
J/gm % J/gm %
Delamination 13.3 68.9 --
+ Bending
Friction + 16.8 87.0 29.5 97.0
Bending
Delamination 2.5 13 0.84 2.8
Bending 10.8 56 18.9 62.2
Friction 6.0 31 10.6 34.8
Comer 0.0 0.0 0.06 0.2
splitting
Total 19.3 100 30.4 100

 

 

 

127

Table 5.14 Carbon composite strip and crush tube energy balance dynamic loading

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

(2.0 m/sec).
Energy Mode Strip Tube
J/ gm % (J/gm) %
Delanrination + 13.0 84.4 --
Bending
Friction + 12.9 83.8 22.9
Bending
Delamination 2.5 16.2 0.84 3.5
Bending 10.5 68.2 18.6 78.1
Friction 2.4 15.6 4.3 18.1
Corner splitting 0.0 0.0 0.06 0.3
Total 15.4 100 23.8 100

 

Table 5.15 Quasi-static vs. dynamic loaded carbon composite tube crush dynamic load

 

 

 

 

 

 

 

 

rate (2.0 m/sec).
Energy Mode Quasi-Static Dynamic Difference
J/ gm % J/gm % J/gm %
Delamination 0.84 2.8 0.84 3.5 0 O
Bendirg 18.9 62.2 18.6 78.1 -0.3 5
Friction 10.6 34.8 4.3 18.1 -6.34 96
Tube Comer 0.063 0.2 0.06 0.3 0
splitting
Total 30.4 100 23.8 100 -6.6 100

 

 

 

 

 

 

 

128

 

 

CHAPTER 6. Conclusions and Future Research

6.1 Introductory Remarks

The need for lightweight materials in automotive applications was introduced in
Chapter 1 with discussion of beneﬁts and challenges of using these materials. One of the
most important challenges to the use of ﬁber-reinforced composite materials for energy
absorbing structures is the understanding of how these materials absorb energy.
Typically, materials to be used for energy absorptions have been characterized using
simple tube structures that are loaded using standard test machines at quasi-static and
dynamic load rates. Simple test tubes are used for testing because of their simple
geometry and convenience. When metallic materials are tested in this way, the speciﬁc
energy absorption, SEA, differs only slightly between quasi-static and dynamic load
rates. Thus, the dynamic behavior of metals can be predicted reasonably well using
computer-aided engineering, CAE, based on quasi-static testing. This approach has had
limited success with composite materials. Testing composite materials at quasi-static
load rate provides SEA values that can be nearly twice as high as dynamic test results.
This observation provides the motive for the research presented in this thesis.

Metallic crush structures absorb energy by plastic deformation. This mode of
energy deformation is consistent with the behavior of the metals used in automotive
energy absorbing structures. Metals used in automotive energy absorbing have a load-
displacement curve that shows a load plateau where the average load changes very little

over a large displacement range. The integrated area under this curve represents energy

absorption.

129

The composite materials tested were observed to absorb energy by several modes.
The energy absorbing modes identiﬁed for the crush tubes considered were: tube corner
splitting, delamination, matrix damage due to bending, and sliding friction. The approach
used to quantify the energy absorption of each mode was to develop simple tests that
would isolate energy absorbing modes.

Mechanical testing of the glass ﬁber and carbon ﬁber composites considered in
this thesis showed elastic behavior. Crush tube comer splitting was measured using a
standard tensile test. Delamination was measured using an end notch ﬂexure test. Matrix
damage due to bending and sliding ﬁiction were measured using a strip test ﬁxture
developed for this purpose.

Tube testing was conducted at a quasi-static load rate for a glass ﬁber/vinyl ester
composite and for a carbon ﬁber/vinyl ester composite. Dynamic testing was performed
at 2.0 m/sec on carbon ﬁber composite tubes. Quasi-static and dynamic testing of carbon
composite coupons was performed at two rates, 0.5 m/sec and 2.0 m/sec, on the carbon
ﬁber composite using the strip test ﬁxture. The higher load rate, 2.0 m/sec, was the
slowest dynamic rate that has been previously tested as shown in Figure 1.3.

6.2 Conclusions

Glass ﬁber/vinyl ester composite crush tubes were investigated with quasi-static
compression and energy absorbing modes were identiﬁed. The observed energy
absorbing modes included tube comer splitting, composite delamination, matrix damage
due to bending, and sliding friction of the composite with the plug type crush trigger.
These same energy-absorbing modes were observed in both quasi-statically compressed

and dynamically crashed carbon composite crush tubes. Energy absorption attributable to

130

comer splitting at quasi-static compression was estimated using standard tensile test
results. Comer splitting was estimated to absorb less that 1% of the total energy absorbed
by both the glass ﬁber composite and the carbon ﬁber composite crush tubes. Energy
absorption attributable to delamination was estimated using the mode H (shear mode)
strain energy release rate obtained using the end notch ﬂexure (ENF) test. Under quasi-
static compression, the glass ﬁber composite delamination SEA was found to be 1.55

J/ gm or 7.5% of the total tube SEA. For carbon ﬁber composite crush tubes, the
delamination SEA was found to be 0.84 J/gm or 2.8%f the total tube SEA.

Experiments seemed to suggest that sliding ﬁiction played an important role in
the energy absorption of composite crush tubes. In an attempt to separate the sliding
friction SEA from the SEA attributable to matrix damage due to bending, an innovative
strip testing ﬁxture was designed, fabricated, and tested and is presented in this thesis.
Carbon ﬁber composite SEA, under quasi-static compression, attributable to matrix
damage due to bending was found to be 18.9 J/gm or 62.2% of the total tube SEA. Under
dynamic crush, the SEA attributable to matrix damage due to bending was found to be
18.6 J/gm or 78.1% of the total tube SEA. Carbon ﬁber composite SEA attributable to
sliding friction under quasi-static compression was found to be 10.6 J/ gm or 34.8% of the
total tube SEA. Under dynamic crush, the SEA attributable to sliding friction was found
to be 4.3 J/gm or 18.1% of the total SEA. The decrease in sliding ﬁiction SEA of 6.3
J/ gm accounted for nearly all of the decrease in tube SEA of 6.6 J/ gm between dynamic
crush and quasi-static compression. Sliding ﬁiction was concluded to be responsible for

the decrease in overall tube SEA from quasi-static compression to dynamic crush.

131

6.3 Future Research

Other composite systems must be evaluated individually until reliable predictive
methods are developed. In the mean time, the physics of damage for each energy
absorption mode identiﬁed should be examined and quantiﬁed at both quasi-static and
dynamic loading rates.

The strip test ﬁxture used in this thesis should be reﬁned so that the strip test
specimens are damaged to a similar level of damage to that observed in tube crush
testing. Possible improvements include guides that eliminate structural instability so that
a thin strip specimen could be compressed. The diameter of roller C, the roller that the
composite strip is forced to bend around, could be reduced to 8 mm so that outer radius of
a 2 mm thick strip specimen is the same as when a 2 mm thick crush tube is forced over

the standard plug trigger.

132

REFERENCES

(1)

(2)

(3)

(4)

(5)

(6)

(7)

(8)

(9)

(10)

(11)

Boeman, R.G. and Caliskan, A.G., "Crush Response of Crash Energy
Management Structures at Intermediate Rates", 2002 Future Car Congress,
Arlington VA, June 2002. (SAE# 2002-01-1954).

Johnson, C. F., "Resin Transfer Molding", Engineering Materials Handbook
Vl-Composites, ASM International 1987

Botkin, M., Fiday, S., and J eryan, R., "Crashworkthiness development of a
Composite Front Structure for a Production Vehicle", Advanced Composites
Conference, Detroit 1997, Engineering Society of Detroit p 11-27

Browne, A. L., Waling, P. 1., Houston, D. Q., Lalik, L., "Generic Tube Crush
Program: Part I: RTM Tubes", 1998 ASME Mechanical Congress and
Exposition, Nov 18, 1998, Anaheim, CA

Hull, Derek, "Energy Absorbing Structural Composites", a report to the
Automotive Composites Consortirun, August, 1994

Automotive Aluminum Crash Energy Management Manual, 1998, The
Aluminum Association, Inc., Washington, D. C.

Thornton, P.H., "Energy Absorption in Composites", Proc SEM Spring
Conference on Experimental Mechanics, June 709, 1993, Dearbom, MI

Farley, Gary, Jones, Robert, "Crushing Characteristics of continuous Fiber-
Reinforced Composite Tubes", Journal of Composite Materials 26 (1992) 27-
50

Farley, Gary, Jones, Robert, "Analogy for the Effect of Material and
Geometrical Variables on Energy-Absorption Capability of Composite
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Mamalis, A. G., Manolakos, D.E., Demosthenous, G. A., Ioannidis, "The
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Mamalis, A. G., Manolakos, D.E., Demosthenous, G. A., Ioannidis, "Analysis
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Elsevier Science Limited, Great Britain

133

(12)

(13)

(14)

(15)

(16)

(17)

(13)

(19)

(20)

(21)
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Chadwick, M. M. and Caliskan, A. G., "Fracture Mechanisms Observed in
Crush of Vinyl Ester-Glass Composite Tubes", Private Communication, 1998

Johnson N. L., Houston, D. Q., Watline, P. J ., Lalic, L., "Generic Tube Crush
Program: Crush Tests of Structural Reaction Injection Molded Tubes",
Crashworthiness, Occupant Protection and Biomechanics in Transportation
Systems-1998, ASME Publication AMD-Vol. 230 BED-Vol. 41, pp 53-74

Bruce, Denise M., Matlock, David K., Speer, John G., De, Amar K.,
"Assessment of the Strain-Rate Dependent Tensile Properties of Automotive
Sheet Steels", Innovations in Steel Sheet and Bar Products and Processing,
and Modeling and Testing of Steel Structures (SP-1837)2004 SAE World
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Weeks, C.A., Sun, C. T., "Modeling Non-Linear Rate-Dependent Behavior in
Fiber-Reinforced Composites", Composites Science and Technology 58
(1998) 603-611, Elsevier Science Limited, Great Britain

Todo, M., Takahashi, K., Beguelin, P., Kausch, H. H., "Strain-rate
dependence on the tensile fracture behavior of woven-cloth reinforced
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771, Elsevier Science Limited, Great Britain

Johnson, N. L., Houston, D. Q., Coyner, R., Lalik, L., "Automotive
Composites Consortium Generic Tube Crush Program: Static Crush Tests of
SRIM Tubes", Proceedings of the 13th Annual ESD Advanced Composites
Conference and Exposition, September 29, 1998, Detroit, MI

Johnson, N. L., "Automotive Composites Consortirun Generic Tube Crush
Program: Dynamic Crush Tests of SRIM Tubes", Proceedings of the 13th
Annual American Society for Composites Conference, September 21-23,
Baltimore, Nﬂ)

Browne, A. L., Houston, D. Q., Lalik, L., "Automotive Composite
Consortium Generic Tube Crush Program: Resin Transfer Molded Tubes",
Crashworthiness, Occupant Protection and Biomechanics in Transportation
Systems-1998, ASME Publication AMD-Vol. 230 BED-Vol.. 41 pp 75-94

Agaram, V., Bilkhu,S., Fidan, S., Botkin, M., and Johnson, N.,"Simulation of
Controlled Failure of Automotive Composite Structures with LS_DYNA3D",
ACCE Conference Proceedings, Dearbom, MI April 7-10, 1997

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ASTM D25 84-02 Standard Test Method for Ignition Loss of Cured
Reinforced Resins

134

(23) Delsen Test Report T 37674, 07/23/03

(24) Ford Central Laboratory Test Report #20679, June 3, 2002, Private
Communication.

135

APPENDIX A

Glass Fiber Composite Crush Tube Load Curves

136

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Carbon Fiber Composite Crush Tube Load Curves

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