Carbon-fiber/Stainless Steel Laminates Improve Composite Bearing Strength

Master of Science Thesis
Development of a
Carbon-ﬁber/Stainless Steel Laminate
Concept
Improving the bearing characteristics of carbon-ﬁber composites by
addition of steel strip reinforcements.
Adam Buczynski
January 27, 2009

Development of a
Carbon-ﬁber/Stainless Steel Laminate
Concept
Improving the bearing characteristics of carbon-ﬁber composites by
addition of steel strip reinforcements.
Master of Science Thesis
For obtaining the degree of Master of Science in Aerospace Engineering
at Delft University of Technology
Adam Buczynski
January 27, 2009
Faculty of Aerospace Engineering · Delft University of Technology

Delft University of Technology
Copyright c Adam Buczynski
All rights reserved.

Delft University Of Technology
Department Of
Aerospace Materials
The undersigned hereby certify that they have read and recommend to the Faculty of
Aerospace Engineering for acceptance a thesis entitled “Development of a Carbon-
ﬁber/Stainless Steel Laminate Concept” by Adam Buczynski in partial fulﬁllment of
the requirements for the degree of Master of Science.
Dated: January 27, 2009
Professor:
prof. dr. ir. R. Benedictus
Supervisor:
dr. ir. R. C. Alderliesten
Readers:
ir. J. Sinke
dr. ir. O. K. Bergsma

Summary
This report presents the development of a Fiber Metal Laminate (FML) concept with stainless
steel and a carbon fiber prepreg as the main constituents. Focus lies on the improvement of
the bearing strength of the carbon fiber composite by the addition of steel sheets and strips.
The goal of this study is to improve the characteristics of carbon fiber composites applied in
joints.
Key questions that have been investigated in this report are:
• Does addition of stainless steel sheets or strips improve the bearing strength of carbon-
fiber composites?
• How does the bearing strength depend on the distribution of the steel within the lami-
nate?
• What are the failure mechanisms involved and are these different from current FMLs?
• Is the addition of steel sheets beneficial despite the high specific weight of steel?
• If so, how much reduction in laminate thickness can be gained by applying the steel
sheets?
Based upon a literature study, the current state of the art of bearing in composites and FMLs
has been discussed, presenting several methods that can be applied to improve the bearing
characteristics of both types of material. Some relevant observations that were reported are:
• Metallic layers usually dominate the bearing strength of FMLs.
• More homogeneous stacking sequences (CP) perform better than less homogeneous ones
(UD), especially for smaller edge distances.
• The favored failure mode for a material undergoing a bearing load is true bearing failure.
M.Sc. thesis Adam Buczynski

vi Summary
• In order to accomplish true bearing failure, sufficiently high values of E/D and W/D
are needed.
• Lateral support in general has a significant favorable effect on the bearing characteristics
of FMLs and CFRPs.
Included in the literature study were also the theoretical descriptions of stresses and strains
in laminates, the related laminate properties and the bearing strength. A discussion of the
production process of current FMLs and of the Carbon-fiber/Stainless Steel laminate concept
has been given to clarify the material selection process, the procedures used to fabricate the
laminate as well as some difficulties that arose during manufacturing of the laminate.
Three series of tests were conducted for this research. First, preliminary tests were performed
to gain insight into what production methods, laminate configurations and materials were
optimal. Second, initial bearing tests were performed with various laminate lay-ups to provide
some initial data on the bearing behavior of the laminate concept. Last, final bearing tests
were carried out to provide enough data to answer the above key questions. The setup, details,
results and observed failure mechanisms of all three test series have been discussed in detail.
From the analysis of the results it was confirmed that regarding laminate configuration, Cross
Ply (CP) performs better than UniDirectional (UD). The effect of the addition of steel strips
on the bearing strength was found to depend greatly on the type of steel and composite
used. Investigation of the distribution of the steel strips within the laminate gives some
contradictory results, but the effect of laminate thickness on the bearing strength follows a
clear trend, improving with thicker panels. A clear benefit of using lateral constraints was
also found.
Based on the analysis of the results, answers to the key questions stated earlier were found.
These answers, which are the main findings of this research, are:
• The addition of thin stainless steel strips does improve the damage bearing strength of
carbon fiber composites. However, this is only the case for certain combinations of steel
and carbon prepreg.
• Positioning the steel in the center of the laminate is better than moving the strips to
the outside. However, this does not seem to be the case when multiple steel sheets are
placed next to each other.
• Bearing failure was the main failure mode occurring during the bearing tests, which is
in conjunction with observations of other FMLs undergoing bearing loads. However,
other failure mechanisms, not previously observed in other FMLs, were also present.
• Despite the high specific weight of steel, the addition of thin strips can be beneficial.
This is particularly true for thicker laminate panels, which have to undergo high bearing
loads. The laminate thickness can potentially be reduced by more than 35% when using
steel strips, without adding more weight to the laminate.
Adam Buczynski M.Sc. thesis

Preface
This report covers my M.Sc. thesis work done at the Delft University of Technology (DUT).
Although at times tough and challenging, it has been an interesting and worthwhile experi-
ence. The work presented in this report would not have been possible without the help of
many people. I would like to mention and thank several of them.
First of all, I would like to thank my supervisor Ren Alderliesten for introducing me to the
project and for assisting me throughout its whole duration. Every meeting with him gave me
a new dose of motivation and inspiration.
Second, I would like to thank my professor Rinze Benedictus for giving me the opportunity
to graduate at the chair of Aerospace Materials. Although it took longer than expected, I
hope the result of my work was worth the wait.
I would further like to thank some of the people who have helped in one way or another with
either the production or the testing of the laminate panels. They are: Michel Badoux, Fred
Bosch, Berthil Grashof, Niels Jalving, Sebastiaan Lindsted, Serge van Meer, Hans Weerheim
and of course Herman Werges.
My thanks also goes out to Esther Rensma and Riccardo Rodi, who have both assisted me
with the manufacturing and curing of my laminate panels.
A part of my work has been performed at the Technion university in Haifa, Israel, which was
a very interesting experience. For this, I would like to express my sincere gratitude to prof.
Haim Abramovic for allowing me to work there and to Anrei Kotler for assisting me with the
equipment and the testing procedures.
Furthermore, I would like to thank my family for their support and faith in me throughout
the project. Last, but not the least, I would like to thank Liat for giving me the motivation
I needed and for her continuous support and encouragement throughout the whole project.
Delft, University of Technology Adam Buczynski
January 27, 2009

viii Preface

Table of Contents
Summary v
Preface vii
List of Figures xiii
List of Tables xv
Acronyms xvii
Nomenclature xvii
1 Introduction 1
1.1 Background of Composites and Fiber Metal Laminates . . . . . . . . . . . . . . 1
1.2 Application of Fiber Metal Laminates . . . . . . . . . . . . . . . . . . . . . . . . 2
1.3 Aim and Contents of this Thesis . . . . . . . . . . . . . . . . . . . . . . . . . . 3
2 Bearing Characteristics of Composites and Fiber Metal Laminates 5
2.1 Introduction . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 5
2.2 Composites versus Fiber Metal Laminates . . . . . . . . . . . . . . . . . . . . . 5
2.3 Optimizing the Bearing Characteristics of Composites and Fiber Metal Laminates 6
2.3.1 Pin Bearing versus Bolt Bearing . . . . . . . . . . . . . . . . . . . . . . 6
2.3.2 Stacking Sequence and Loading Direction . . . . . . . . . . . . . . . . . 6
2.3.3 Specimen Dimensions . . . . . . . . . . . . . . . . . . . . . . . . . . . . 6
2.3.4 Bolt Hole Clearance . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 7
2.3.5 Fiber Steering and Matrix Stiﬀening . . . . . . . . . . . . . . . . . . . . 7
2.3.6 Local Reinforcements . . . . . . . . . . . . . . . . . . . . . . . . . . . . 7
2.4 Failure modes and fracture mechanisms . . . . . . . . . . . . . . . . . . . . . . 7
2.4.1 Bearing Failure . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 8
2.4.2 The Eﬀect of Lateral Restraints . . . . . . . . . . . . . . . . . . . . . . 9
2.4.3 Fracture Mechanisms . . . . . . . . . . . . . . . . . . . . . . . . . . . . 10
2.5 Summary . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 10

x Table of Contents
3 Theory and Definitions 11
3.1 Classical Laminate Theory . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 11
3.1.1 Basic Assumptions . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 11
3.1.2 Stresses and strains per layer . . . . . . . . . . . . . . . . . . . . . . . . 12
3.1.3 Stresses and Strains in the Complete Laminate . . . . . . . . . . . . . . 13
3.1.4 Curing Stresses . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 14
3.1.5 External Stresses . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 14
3.1.6 Summary . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 15
3.2 Basic Mechanical Laminate Properties . . . . . . . . . . . . . . . . . . . . . . . 15
3.2.1 The Metal Volume Fraction Method . . . . . . . . . . . . . . . . . . . . 15
3.2.2 Modulus of Elasticity and the Shear Modulus . . . . . . . . . . . . . . . 15
3.2.3 Tensile Yield Strength and Ultimate Strength . . . . . . . . . . . . . . . 16
3.2.4 Bearing Yield Strength and Ultimate Strength . . . . . . . . . . . . . . . 16
4 Manufacturing of the Laminate 17
4.1 Introduction . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 17
4.2 Manufacturing of Current Fiber Metal Laminates . . . . . . . . . . . . . . . . . 17
4.2.1 Manufacturing of GLARE . . . . . . . . . . . . . . . . . . . . . . . . . . 17
4.2.2 Manufacturing of ARALL . . . . . . . . . . . . . . . . . . . . . . . . . . 18
4.2.3 Manufacturing of CARE . . . . . . . . . . . . . . . . . . . . . . . . . . 18
Galvanic corrosion . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 18
Pre-treatment and curing . . . . . . . . . . . . . . . . . . . . . . . . . . 19
4.2.4 Manufacturing of TiGr . . . . . . . . . . . . . . . . . . . . . . . . . . . 19
4.3 Manufacturing of the Carbon-Fiber/Stainless Steel Laminate . . . . . . . . . . . 19
4.3.1 Choice of Materials and their Properties . . . . . . . . . . . . . . . . . . 19
Stainless steel . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 19
Carbon fiber/epoxy . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 20
4.3.2 Pre-treatment of the Steel Sheets . . . . . . . . . . . . . . . . . . . . . 20
4.3.3 Lay-up of the Laminate Panels . . . . . . . . . . . . . . . . . . . . . . . 21
4.3.4 The Curing Cycle . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 21
4.3.5 Testing for Inclusions and Contaminations . . . . . . . . . . . . . . . . . 22
4.3.6 Production of the Test Specimens . . . . . . . . . . . . . . . . . . . . . 22
4.4 Concluding Remarks . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 22
5 Testing the Laminate Concept 23
5.1 Introduction . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 23
5.2 Preliminary Tensile Tests . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 23
5.2.1 Laminate Configuration . . . . . . . . . . . . . . . . . . . . . . . . . . . 24
5.2.2 Specimen Geometry . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 24
5.2.3 Testing Conditions and Procedure . . . . . . . . . . . . . . . . . . . . . 24

Table of Contents xi
5.2.4 Test Results . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 24
5.2.5 Brief Discussion of the Results . . . . . . . . . . . . . . . . . . . . . . . 25
5.3 Initial Bearing Tests . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 25
5.3.3 Test Conditions and Procedure . . . . . . . . . . . . . . . . . . . . . . . 26
5.3.4 Test Results . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 26
5.3.5 Failure Modes . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 28
Observations . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 28
Exceptions . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 29
Relation of the failure mode to the laminate lay-up . . . . . . . . . . . . 30
5.4 Final Bearing Tests . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 30
5.4.3 Test Conditions and Procedure . . . . . . . . . . . . . . . . . . . . . . . 32
5.4.4 Test Results . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 32
5.4.5 Failure Modes . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 33
Observations . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 33
Exceptions . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 36
6 Discussion of the Test Results 39
6.1 Effect of the Laminate Configuration . . . . . . . . . . . . . . . . . . . . . . . . 39
6.2 Effect of the Addition of Steel . . . . . . . . . . . . . . . . . . . . . . . . . . . 39
6.3 Effect of the Distribution of the Steel Strips . . . . . . . . . . . . . . . . . . . . 42
6.4 Effect of Laminate Thickness . . . . . . . . . . . . . . . . . . . . . . . . . . . . 44
6.5 Effect of Providing Lateral Constraint . . . . . . . . . . . . . . . . . . . . . . . 46
6.6 Comparison of the Initial and Final Bearing Tests . . . . . . . . . . . . . . . . . 47
6.6.1 Comparison of the Test Results . . . . . . . . . . . . . . . . . . . . . . . 47
6.6.2 Estimate of Laminate Performance for a Different Material Combination . 48
6.6.3 Comparison of the Failure Mechanisms . . . . . . . . . . . . . . . . . . . 49
7 Conclusions and Recommendations 51
7.1 Conclusions . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 51
7.2 Recommendations . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 52
A Preliminary Test Results 59
B Initial Bearing Test Results 61
C Final Bearing Test Results 65

xii Table of Contents

List of Figures
2.1 Typical bearing failure modes[21] . . . . . . . . . . . . . . . . . . . . . . . . . . 8
2.2 Illustration of the damage bearing load and the failure bearing load[14] . . . . . . 9
5.1 Typical bearing test curve . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 27
5.2 Typical failure stages during bearing failure (not the same specimen) . . . . . . . 28
5.3 Crack formation at the top of the specimens (not the same specimen) . . . . . . 29
5.4 Close-up cross-section view of a reinforced specimen . . . . . . . . . . . . . . . . 29
5.5 Failure mode exceptions . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 30
5.6 Close-up cross-section view showing buckling of the inner sheets . . . . . . . . . 31
5.7 Typical bearing test curve (with lateral constraint) . . . . . . . . . . . . . . . . . 33
5.8 Typical bearing test curve (without lateral constraint) . . . . . . . . . . . . . . . 34
5.9 Typical observed failure modes (unconstrained specimens) . . . . . . . . . . . . 35
5.10 Typical observed failure modes (laterally constrained specimens) . . . . . . . . . 36
5.11 Typical observed failure modes (composite-only panel) . . . . . . . . . . . . . . 36
5.12 Specimen damage after large displacements . . . . . . . . . . . . . . . . . . . . 37
5.13 Cross-section views showing outward bending of steel layers . . . . . . . . . . . . 37
6.1 The effect of the addition of steel strips on the damage bearing strength . . . . . 40
6.2 Relative damage bearing load of all the panels of the final bearing tests . . . . . 41
6.3 Metal volume fraction plotted against the relative damage bearing load . . . . . 41
6.4 The effect of the distribution of the steel strips on the damage bearing strength . 42
6.5 Cross-section view of P3A specimens . . . . . . . . . . . . . . . . . . . . . . . . 43
6.6 The effect of double metal sheets on the damage bearing strength . . . . . . . . 43
6.7 Cross-section view of P3C specimens . . . . . . . . . . . . . . . . . . . . . . . . 44
6.8 The effect of laminate thickness on the damage bearing strength . . . . . . . . . 45

xiv List of Figures
6.9 Thickness of the laminate vs. the damage bearing load . . . . . . . . . . . . . . 45
6.10 Thickness vs. the weight of the laminate with constant damage bearing load lines 46
6.11 Cross-section view of P3D specimens . . . . . . . . . . . . . . . . . . . . . . . . 46
6.12 Behavior of the load-displacement curves after the damage bearing load . . . . . 47
6.13 Metal volume fraction versus estimated relative damage bearing load . . . . . . . 48
6.14 Thickness vs. weight with estimated damage bearing load lines . . . . . . . . . . 49
A.1 Tensile Test Results of Panel P1A . . . . . . . . . . . . . . . . . . . . . . . . . 59
A.2 Tensile Test Results of Panel P1B . . . . . . . . . . . . . . . . . . . . . . . . . 60
B.1 Bearing Test Results of Panel P2E-NR . . . . . . . . . . . . . . . . . . . . . . . 61
B.2 Bearing Test Results of Panel P2B-NR . . . . . . . . . . . . . . . . . . . . . . . 62
B.3 Bearing Test Results of Panel P2E-R2C . . . . . . . . . . . . . . . . . . . . . . 62
B.4 Bearing Test Results of Panel P2C-R2C . . . . . . . . . . . . . . . . . . . . . . 63
B.5 Bearing Test Results of Panel P2C-R2G . . . . . . . . . . . . . . . . . . . . . . 63
B.6 Bearing Test Results of Panel P2B-R1M . . . . . . . . . . . . . . . . . . . . . . 64
C.1 Bearing Test Results of Panel P3A-I . . . . . . . . . . . . . . . . . . . . . . . . 65
C.2 Bearing Test Results of Panel P3A-II . . . . . . . . . . . . . . . . . . . . . . . . 66
C.3 Bearing Test Results of Panel P3A-III . . . . . . . . . . . . . . . . . . . . . . . 66
C.4 Bearing Test Results of Panel P3A-IV . . . . . . . . . . . . . . . . . . . . . . . 67
C.5 Bearing Test Results of Panel P3A-V . . . . . . . . . . . . . . . . . . . . . . . . 67
C.6 Bearing Test Results of Panel P3B (unconstrained) . . . . . . . . . . . . . . . . 68
C.7 Bearing Test Results of Panel P3B (lateral constraint) . . . . . . . . . . . . . . 68
C.8 Bearing Test Results of Panel P3C-I . . . . . . . . . . . . . . . . . . . . . . . . 69
C.9 Bearing Test Results of Panel P3C-II . . . . . . . . . . . . . . . . . . . . . . . . 69
C.10 Bearing Test Results of Panel P3C-III . . . . . . . . . . . . . . . . . . . . . . . 70
C.11 Bearing Test Results of Panel P3C-IV . . . . . . . . . . . . . . . . . . . . . . . 70
C.12 Bearing Test Results of Panel P3D-I . . . . . . . . . . . . . . . . . . . . . . . . 71
C.13 Bearing Test Results of Panel P3D-II . . . . . . . . . . . . . . . . . . . . . . . . 71
C.14 Bearing Test Results of Panel P3D-III . . . . . . . . . . . . . . . . . . . . . . . 72
C.15 Bearing Test Results of Panel P3E . . . . . . . . . . . . . . . . . . . . . . . . . 72

List of Tables
2.1 Optimal E/D and W/D ratios for CFRPs, GFRPs and certain FMLs . . . . . . . 7
4.1 Overview of material properties . . . . . . . . . . . . . . . . . . . . . . . . . . . 20
5.1 Preliminary Tensile Test Laminate Configurations . . . . . . . . . . . . . . . . . 24
5.2 Tensile Test Results . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 24
5.3 Initial Bearing Test Laminate Configurations . . . . . . . . . . . . . . . . . . . . 26
5.4 Initial Bearing Test Results (Average Values) . . . . . . . . . . . . . . . . . . . 27
5.5 Final Bearing Test Laminate Configurations . . . . . . . . . . . . . . . . . . . . 31
5.6 Final Bearing Test Results (Average Values) . . . . . . . . . . . . . . . . . . . . 35

xvi List of Tables

Acronyms
ARALL Aramid Reinforced ALuminium Laminate
BL Bearing Load
CAA Chromic Acid Anodizing
CARALL CArbon fiber Reinforced ALuminium Laminate
CARE CArbon fiber REinforced laminate
CFRP Carbon Fiber Reinforced Plastic
CLT Classical Laminate Theory
CP Cross Ply
DBL Damage Bearing Load
DUT Delft University of Technology
FML Fiber Metal Laminate
FRP Fiber Reinforced Plastic
GLARE GLAss fiber REinforced laminate
GFRP Glass Fiber Reinforced Plastic
MVF Metal Volume Fraction
PAA Phosphoric Acid Anodizing
PEI PolyEtherImide
SHA Sodium Hydroxide Anodizing
TiGr Titanium Graphite laminate
UD UniDirectional

xviii Acronyms

Chapter 1
Introduction
1.1 Background of Composites and Fiber Metal Laminates
Although the concept of composite materials with a fibrous reinforcement is very old and
dates back thousands of years, it has not been until recently that this concept found its way
into the aircraft industry. The horizontal stabilizer of the F-111, developed in 1964, was
the first primary aircraft structure made of fiber reinforced materials. Usage continued in the
1970s with the F-18 Hornet and the AV-8B Harrier, the latter having its entire wing structure
made of a carbon fiber-epoxy composite[1].
During the last few decades, a lot of research was performed to develop an optimized laminated
composite material for application in fatigue sensitive areas of modern civil aircraft. This
research eventually resulted in the development of the Fiber Metal Laminate (FML) concept
at the Delft University of Technology (DUT). FMLs consist of metal sheets bonded together
with alternating layers of fibers impregnated in an epoxy matrix. The advantages of the
high strength isotropic metal sheets with fracture resistant fibers are thus combined, leading
to substantial weight savings and superior fatigue and static properties. The crack-bridging
property of the fibers produces an effective decrease or even a stop in (fatigue) crack growth
occurring in the metal sheets[2].
Two variants of the FML concept are aluminium sheets reinforced with aramid fibers or
aluminium sheets reinforced with glass fibers, now commercially available as respectively
ARALL and GLARE. The latter is currently being used in the upper skin of the Airbus
A-380 fuselage. Both ARALL and GLARE have shown to exhibit excellent fatigue crack
propagation behavior.
Some studies of the usage of carbon fibers in an FML concept have also been performed.
Both the CARE and CARALL[3] laminates use aluminium sheets with carbon fibers as a
reinforcement. Although these laminates have demonstrated a high strength and good fatigue
behavior, the combination of carbon fibers with aluminium sheets is problematic due to the
occurrence of galvanic corrosion between these two constituents[4,5].

2 Introduction
Titanium sheets in combination with carbon fibers have also been investigated, which resulted
in the laminate called TiGr. Although good for high temperature applications, there are some
drawbacks to the usage of titanium. When compared to aluminium, material cost for is higher,
machinability is harder and the pre-treatment for laminate bonding is more difficult[4].
The usage of steel sheets in combination with carbon fibers in FML concepts has been in-
vestigated briefly, both as a stand-alone laminate[1,7] and combined with aluminum and glass
fiber layers[8,9]. In addition, the usage of steel reinforcement strips in GLARE has also been
investigated[4]. But neither of these concepts have been developed into an application yet. It
is believed however that this combination of materials has a great potential in future aircraft
structures. Thin steel strips added to a carbon-fiber composite laminate as a local reinforce-
ment, or steel sheets to support the laminate as a whole in an FML concept, could prove to
be beneficial despite the high specific weight of steel.
1.2 Application of Fiber Metal Laminates
Due to their good mechanical properties, FMLs can replace aluminium alloys in some major
parts of an aircraft structure. Structures which will benefit most from these FMLs are struc-
tures for which fatigue and damage tolerance are important design criteria. Examples of these
kind of structures are the fuselage, the lower wing skin panels or the horizontal stabilizer.
Potential applications of FMLs consisting of stainless steel and carbon fibers could be highly
fatigue loaded lugs, strap material for selective reinforced aluminium structures, or fuselage
and wing skin applications.
When designing aircraft structures, static strength properties like tension, compression and
shear play an important role. However, due to the limited width of metal sheets, the necessity
to join parts of different materials together as well as the requirement of easy assembly and
disassembly, mechanical joints (rivetted or bolted) are often required in aircraft structures.
Often such joints are the weakest points in a structure. The need to design them correctly
means another strength parameter comes into play, namely the bearing strength of a material.
The bearing strength can be seen as the resistance of a material to a fastener loaded hole.
This resistance depends to a large extent on a parameter called the edge distance of the
hole. This is the distance between the center of the hole and the edge of the material sheet in
loading direction. Designing for minimum weight implies that this distance should be as small
as possible. This makes the bearing strength a critical design parameter. Fiber reinforced
materials in particular have a rather low resistance to fastener loads due to their anisotropy
and low shear resistance. Relatively large and undesirable edge distances are therefore often
required when using such materials (like FMLs) in mechanical joints[4,6].
Another important parameter is the blunt notch strength of a material. Due to the presence
of fastener holes, stress concentrations arise within the structure. These stress concentrations
could lead to premature failure at load levels below the ultimate strength of the material.
In general, blunt notches are more critical for fiber reinforced materials and FMLs than for
metals, due to the brittle nature of the fibers and the anisotropic characteristics of the fiber
layer. This can result in increased stress concentrations around the notch[4].

1.3 Aim and Contents of this Thesis 3
1.3 Aim and Contents of this Thesis
The aim of this thesis will be to develop a Carbon-fiber/Stainless Steel laminate concept, using
stainless steel sheets or strips in order to improve the bearing characteristics of a carbon-fiber
composite laminate.
The research and tests performed in this thesis will try to answer several key questions related
to this topic:
• Does addition of stainless steel sheets or strips improve the bearing strength of carbon-
fiber composites?
• How does the bearing strength depend on the distribution of the steel within the lami-
nate?
• What are the failure mechanisms involved and are these different from current FMLs?
• Is the addition of steel sheets beneficial despite the high specific weight of steel?
• If so, how much reduction in laminate thickness can be gained by applying the steel
sheets?
Chapter 2 will present the current state of the art of bearing in composites and FMLs.
The latest findings and developments will be presented, common failure modes and fracture
mechanisms will be analyzed, and methods found to improve the bearing strength will be
discussed.
In Chapter 3 the Classical Laminate Theory will be introduced as well as some derivations
for the application of this theory. Some basic definitions used further in the thesis will also
be discussed briefly.
Next, a discussion of the production process of current FMLs and of the Carbon-
fiber/Stainless Steel laminate will be presented in Chapter 4. The procedures that were
used to fabricate the laminate will be discussed in detail.
Three series of tests were conducted for this research. First, preliminary tests were performed
to gain insight into what production methods, laminate configurations and materials were
optimal. Second, initial bearing tests were performed with various laminate lay-ups to provide
some initial data on the bearing behavior of the laminate concept. Last, final bearing tests
were carried out to provide enough data to answer the above questions. Chapter 5 will present
the setup, details and results of all three of these test series. Observed failure mechanisms
will be discussed as well.
In Chapter 6 the results of the initial and final bearing tests will be discussed in detail. This
chapter will also try to answer the questions posed in this section.
The conclusions of this report and any recommendations for future work will be presented in
Chapter 7.

4 Introduction

Chapter 2
Bearing Characteristics of Composites
and Fiber Metal Laminates
This chapter gives an overview of the latest findings on the bearing characteristics of both
composites and FMLs. Currently known and researched methods to improve these charac-
teristics are presented as well.
2.1 Introduction
Although the fiber content in FMLs has resulted in improved damage tolerance and fatigue
properties compared to monolithic metal sheets, fiber addition also negatively influences sev-
eral static properties.
Biaxial (Cross Ply) laminates in particular suffer from a decreased axial stiffness both in
tension and in compression. This lower stiffness results in a lower yield strength compared
to the metal sheet. Further, the bearing strength of unidirectional fiber layers is low, which
negatively influences the strength of mechanical joints. For small edge distances in particular,
the fiber layers cannot reach their complete strength and this might result in premature failure.
An increase of notch sensitivity for both blunt and sharp notches is found as well. Holes and
cracks in the material therefore become a more critical parameter for the design of FMLs[4,10].
2.2 Composites versus Fiber Metal Laminates
When considering the bearing characteristics of Fiber Reinforced Plastics (FRP) versus FMLs,
one finds that the presence of one or more metallic layers generally benefits laminates in
terms of bearing strength. In FMLs it is the metal layers that dominate the bearing yield
and ultimate strength. The fiber layer contribution is generally limited, due to its relatively
low shear properties[10].

6 Bearing Characteristics of Composites and Fiber Metal Laminates
2.3 Optimizing the Bearing Characteristics of Composites and
Fiber Metal Laminates
A lot of research has been done to study and improve the behavior of composites and FMLs
undergoing bearing loads. In this section an overview of this research will be given and various
approaches that can be taken to improve the bearing characteristics of FMLs will be outlined.
2.3.1 Pin Bearing versus Bolt Bearing
In general there are two distinct test methods possible when testing the bearing characteristics
of composites or Fiber Metal Laminates. They are pin bearing, with the specimen loaded
by a pin, and bolt bearing, with the specimen being clamped and loaded by a bolt. In the
case of pin bearing, it is possible to perform the bearing tests with or without added lateral
support. It has been observed that finger-tight lateral support limits the amount of layer
buckling and delamination that occurs during loading, which results in considerably better
bearing performance than for cases lacking lateral support[11,12].
In the case of FMLs, it has been observed that in a pin bearing configuration delamination
buckling of the aluminum layers precedes joint collapse and that this phenomenon is thus
responsible for failure of the joint. Observations made during optical microscopy after testing
the laminates revealed extensive delamination and buckling of the aluminum and fiber layers,
which further supports this notion[11,13,14].
2.3.2 Stacking Sequence and Loading Direction
Early tests on Carbon Fiber Reinforced Plastics (CFRP) showed that the influence of stacking
sequence of quasi isotropic laminates on the bearing strength is minor, but that the stacking
sequence does have an effect on local failure mechanisms[15]. Later tests however showed that
the bearing strength was up to 12% higher for quasi isotropic CFRPs where the longitudinal
fiber layers were located in the center, as opposed to the sides of the laminate[16,17]. Interest-
ingly enough, the opposite was found to be true in the case of Glass Fiber Reinforced Plastics
(GFRP)[18].
Less homogeneous stacking sequences, for example UniDirectional (UD) compared to Cross
Ply (CP) configurations, were found to exhibit lower bearing strengths. This effect seems to
be particularly pronounced for relatively small edge distances[18].
Tests on various configurations of GLARE indicated that there was no profound effect of
loading direction on the bearing strength of this material. The bearing strength of FMLs does
decrease at elevated temperatures, but this effect is less pronounced when the Metal Volume
Fraction (MVF) of the laminate is higher. This implies that the bearing characteristics of an
FML are dominated by the behavior of the metal layers[4,19].
2.3.3 Specimen Dimensions
Many researchers investigated the effect of the width-to-diameter (W/D) and edge-distance-
to-diameter (E/D) ratios on the bearing strength of both CFRPs, GFRPs and FMLs. These

2.4 Failure modes and fracture mechanisms 7
ratios need to be sufficiently high for true bearing to be achieved. Lower than optimal
values for W/D or E/D can lead to (undesirable) net tension or shear-out failure modes
respectively[14]. Section 2.4 deals with the differences between these failure modes in more
detail.
A summary of optimal W/D and E/D ratios is given in Table 2.1. The values for these ratio
differ slightly depending on the type of material and on the stacking sequence in some cases.
Material Optimal E/D ratio Optimal W/D ratio
CFRP[16,17,20] ≥ 4 ≥ 4
GFRP[18] ≥ 3 3 ≤ x ≤ 4
GLARE[14,6] ≥ 2, ≥ 2.5 ≥ 2
TiGr[21] 2 ≤ x ≤ 3 -
Table 2.1: Optimal E/D and W/D ratios for CFRPs, GFRPs and certain FMLs
2.3.4 Bolt Hole Clearance
A negative effect of bolt hole clearance on the bearing strength of CFRPs was found at 4%
hole deformation, significantly reducing the bearing strength of the joints. However, the
ultimate bearing strength of the joints does not seem to depend significantly on the bolt hole
clearance[22].
2.3.5 Fiber Steering and Matrix Stiffening
Recently several other, less conventional methods to improve the bearing characteristics of
CFRPs have also been investigated. It was found that the bearing strength could be improved
by up to 35% by means of fiber steering. Stiffening of the matrix by using clay nanoparticles
was also attempted, but this did not result in an improved bearing strength due to the
occurrence of a premature, unspecified failure mode. The incorporation of the nanoparticles
did however stiffen the bearing response, indicating that the method could produce improved
bearing strength if introduction of the new and premature failure mode could be avoided
somehow[23,24].
2.3.6 Local Reinforcements
Less research has been done on the addition of local reinforcements to a laminate lay-up
in order to improve bearing characteristics. One particular study, focussed on the bearing
characteristics of GLARE, showed however that the addition of stainless steel strip reinforce-
ments in between the layers of the laminate could be used to improve the bearing strength of
FMLs[4].
2.4 Failure modes and fracture mechanisms
There are four basic failure modes in which a bolted joint undergoing tension can fail. These
are:

• net failure of the section
• shear-out of the bolt (cleavage failure)
• bearing failure
• transverse splitting (delamination)
Three of these failure modes are illustrated in Figure 2.1. It has been determined that the
kind of failure mode that will occur, often depends on the aforementioned width-to-diameter
(W/D) and edge-distance-to-diameter (E/D) ratios. Low values for the W/D ratio result in
net tension failure, whereas an insufficient E/D ratio results in shear-out failure[14].
Figure 2.1: Typical bearing failure modes[21]
Out of the four failure modes mentioned, bearing failure is by far the most preferable mode,
because in this mode the joined members are not abruptly and catastrophically separated. In-
stead, damage growth is stable and the joint will generally still be able to withstand significant
loads after the first failure[12,14,25].
The following subsections will therefore deal primarily with the bearing failure mode and its
typical properties and fracture mechanisms.
2.4.1 Bearing Failure
Bearing failure of bolted or pin loaded composite or FML joints is often a complicated process
in which many parameters are involved. A reason for this complexity is the out-of-plane
compressive deformation that occurs in the vicinity of the hole. Characteristic features of
bearing failure include[11,26]:
• elongation of the fastener hole
• metal layer macro buckling
• fiber layer micro buckling
• matrix cracking
• delamination and yielding of the metal layers
• out-of-plane shear cracking

2.4 Failure modes and fracture mechanisms 9
An extensive experimental investigation[26] was performed on the bearing strength and failure
behavior of bolted composite joints. Based on the results it was concluded that bearing failure
in composites is a process of compressive damage accumulation, where four macroscopic stages
can be identified:
• damage onset
• damage growth
• local fracture
• structural fracture
In the case of FMLs, similar stages can be observed. However, the dominant bearing fail-
ure features seem to be buckling and delamination of the metal layers, which precede joint
collapse[12].
2.4.2 The Effect of Lateral Restraints
Different behavior is exhibited by FMLs undergoing bearing loading depending on whether
or not there are lateral restraints present.
Typical load-displacement curves for pin loaded joints without lateral support show a single
maximum value, which coincides with the first failure of the specimen. Laterally constrained
or bolt loaded curves however, exhibit two or more maxima with the second maximum (failure
bearing load) invariably higher than the first maximum (damage bearing load). This behavior
can be observed in Figure 2.2.
Figure 2.2: Illustration of the damage bearing load and the failure bearing load[14]
Bolted and laterally restricted joints have shown significantly larger bearing strengths than
pin loaded joints without lateral support. The amount of clamping force applied on the bolt
however, was found to hardly influence the damage bearing strength. It does seem to have a
favorable influence on the maximum failure bearing strength[14].

2.4.3 Fracture Mechanisms
Two shear buckling phenomena were observed in bearing failure of FMLs after microscopic
examinations: one occurring just outside the washers area, and the other involving the ma-
terial directly constrained by the washers. Along the entire buckled length of the material,
delamination was observed as well. From these observations the following sequence of failure
events was hypothesized[14]:
• delamination inducing buckling just outside the washers,
• buckling in these layers in shear mode,
• buckling in the layers between the washers
This sequence might explain why there is hardly any effect of the clamping force on the
damage bearing strength.
2.5 Summary
Since the current study focusses on an FML type of material, the following observations are
of relevance:
• Metallic layers usually dominate the bearing strength of FMLs.
• More homogeneous stacking sequences (CP) perform better than less homogeneous ones
(UD), especially for smaller edge distances.
• The favored failure mode for a material undergoing a bearing load is true bearing failure.
• In order to accomplish true bearing failure, sufficiently high values of E/D and W/D
are needed.
• Lateral support in general has a significant favorable effect on the bearing characteristics
of FMLs and CFRPs.

Chapter 3
Theory and Definitions
This chapter will briefly go over the basics of Classical Laminate Theory (CLT) and it will
cover methods for the determination of elastic properties of FMLs. The chapter will also
cover all the definitions that are used further on in this report.
3.1 Classical Laminate Theory
This section discusses the derivation of the CLT for FML’s and will state the assumptions
under which the theory is valid. This derivation has been summarized by Homan in[27] and
is included here for completeness.
3.1.1 Basic Assumptions
The following assumptions are used in the CLT[28]:
1. Each layer (lamina) of the laminate is quasihomogeneous and orthotropic.
2. The laminate is thin, with its lateral dimensions much larger than its thickness, and is
loaded in its plane only. Thus, the laminate and its layers (except for their edges) are
in a state of plane stress (σz = τxz = τyz = 0).
3. All displacements are small compared to the thickness of the laminate (|u|, |v|, |w| h).
4. Displacements are continuous throughout the laminate.
5. In-plane displacements vary linearly through the thickness of the laminate. Thus, u and
v displacements in the x− and y−directions are linear functions of z.
6. Transverse shear strains γxz and γyz are negligible. This assumption and the preceding
one imply that straight lines normal to the middle surface remain straight and normal
to that surface after deformation.

12 Theory and Definitions
7. Strain-displacement and stress-strain relations are linear.
8. Normal distances from the middle surface remain constant. Thus, the transverse normal
strain z is negligible compared to the in-plane strains x and y.
3.1.2 Stresses and strains per layer
The stresses and strains in each layer of the FML are related by the generalized Hooke’s law.
This law states:
¯σ = S¯ (3.1)
¯ = C¯σ (3.2)
where:
¯σ =


σx
σy
τxy

 (3.3)
¯ =


x
y
γxy

 (3.4)
The coordinates x and y coincide with the materials principal axes. The compliance matrix
C can be written as:
C =



1
Ex
−νxy
Ex
0
1
Ey
0
1
Gxy


 (3.5)
and the stiffness matrix S can be written as:
S =



Ex
1−νxyνyx
νxyEyνxy
1−νxyνyx
0
Ey
1−νxyνyx
0
Gxy


 (3.6)
The stiffness properties under an angle φ are:
¯σφ = M ¯σ (3.7)
¯ = MT
¯φ (3.8)

3.1 Classical Laminate Theory 13
With M being the off-axis matrix:
M =


cos2 φ sin2
φ 2 cos φ sin φ
sin2
φ cos2 φ −2 cos φ sin φ
− cos φ sin φ cos φ sin φ cos2 φ − sin2
φ

 (3.9)
The stiffness and compliance matrix for laminates under an angle φ with respect to the
material principle axes follow from:
¯σφ = MSMT
¯φ = Sφ¯φ, (3.10)
Sφ = MSMT
¯φ = [M−1
]T
CM−1
¯σφ = Cφ¯σφ, (3.11)
Cφ = [M−1
]T
CM−1
The inverse matrix of M can be written as:
M−1
=


cos2 φ sin2
φ −2 cos φ sin φ
sin2
φ cos2 φ 2 cos φ sin φ
cos φ sin φ − cos φ sin φ cos2 φ − sin2
φ

 (3.12)
3.1.3 Stresses and Strains in the Complete Laminate
With the stresses and strains per layer known, the properties for a layer p can be written as:
(¯σφ)p = (Sφ)p¯φ (3.13)
Standard FML grades are defined such that the angle φ is the same for all layers. The
properties for n layers can then be obtained by:
(¯σφ)lam =
n
p=1
(¯σφ)p
tp
tlam
=
n
p=1
(Sφ)p
tp
tlam
¯φ = (Sφ)lam¯φ (3.14)
The stiffness and compliance matrices for the laminate can then be written as:
(Sφ)lam =
n
p=1
(Sφ)p
tp
tlam
(3.15)
(Cφ)lam =
n
p=1
(Cφ)p
tp
tlam
(3.16)
(3.17)

3.1.4 Curing Stresses
Cooling down from curing temperature will cause a strain in the laminate, because the metal
and prepreg layers will have different coefficients of thermal expansion. However, because all
of the layers are attached to each other, all of them must comply with this strain. This leads
to the following equilibrium:
n
p=1
(Sφ)ptp ¯αp = (Sφ)lamtlam ¯αlam (3.18)
with ¯αp being the vector with the thermal expansion coefficient for layer p:
¯αp =


αx cos(φ) + αy sin(φ)
αx sin(φ) + αy cos(φ)
0

 (3.19)
And for the laminate:
¯αlam =
1
tlam
(Sφ)−1
lam
n
p=1
(Sφ)ptp ¯αp (3.20)
The strain due to the thermal expansion is:
¯cure = ¯αlam∆T (3.21)
where:
∆T = Tenv − Tcure (3.22)
Then the internal stresses per layer due to curing follow as:
¯σcure,p = (Sφ)p(¯cure − ∆T ¯αp) (3.23)
3.1.5 External Stresses
Assuming an external stress σ acting on the laminate:
¯σ = (¯σ)lam (3.24)
The strain can then be written as:
¯ = (Sφ)−1
lam(¯σ)lam (3.25)
The stress level per layer due to the external stress will then be:
(¯σφ)p = (Sφ)p¯ = (Sφ)p(Sφ)−1
lam(¯σ)lam (3.26)

3.2 Basic Mechanical Laminate Properties 15
3.1.6 Summary
The total (or gross) stress lever in a layer p is the sum of the curing stress and the stress due
to external loading. This stress level can be written as:
(¯σφ)p = (Sφ)p (Sφ)−1
lam(¯σ)lam + ∆T(αlam − αp) (3.27)
The stiffness matrices can be derived from Section 3.1.2 for layer p and from Section 3.1.3 for
the complete laminate. The thermal expansion coefficients can be derived from Section 3.1.4.
The values (¯σ)lam and ∆T are input parameters.
3.2 Basic Mechanical Laminate Properties
The overall behavior of a laminate is a function of the properties and stacking sequence of
its individual layers. The CLT as derived above in Section 3.1 predicts this behavior of the
laminate within its framework of assumptions.
Some basic mechanical laminate properties in longitudinal direction however can also be
reasonably accurately predicted using the so called Metal Volume Fraction.
3.2.1 The Metal Volume Fraction Method
The Metal Volume Fraction (MVF) of an arbitrary elastic-plastic FML is defined as:
MV F =
nm · tm
nm · tm + npr · tpr
(3.28)
with tm and tpr the thickness of the metal layers and prepreg layers respectively, and nm and
npr the amount of metal and prepreg layers.
Using the MVF, several laminate properties can be predicted with reasonable accuracy. Meth-
ods to do this are presented below[1,4,29].
3.2.2 Modulus of Elasticity and the Shear Modulus
The modulus of elasticity in longitudinal and latitudinal directions, as well as the shear
modulus can be predicted as follows:
Ex,lam = MV F · Em + (1 − MV F)Ex,pr (3.29)
Ey,lam = MV F · Em + (1 − MV F)Ey,pr (3.30)
Gxy,lam = MV F · Gxy,m + (1 − MV F)Gxy,pr (3.31)
(3.32)
with Em the metal layer modulus, Ex,pr the prepreg layer longitudinal modulus, Ey,pr the
prepreg layer latitudinal modulus, Gxy,m the metal layer shear modulus and Gxy,pr the prepreg
layer shear modulus respectively[30].

3.2.3 Tensile Yield Strength and Ultimate Strength
The laminate yield strength and ultimate strength in tensile direction can be predicted with
the following two equations:
σ0.2,lam = MV F + (1 − MV F)
Ex,pr
Em
σ0.2,m (3.33)
σult,lam = MV F · σult,m + (1 − MV F)σult,pr (3.34)
with σ0.2,m the metal layer yield strength, σult,m the metal layer ultimate strength and σult,pr
the prepreg layer ultimate strength. The prediction of σult,lam is quite crude however, since
it assumes simultaneous failure of all components of the laminate. In general, this does not
happen.
3.2.4 Bearing Yield Strength and Ultimate Strength
The bearing yield strength (2% permanent hole deformation) and bearing ultimate strength
of the laminate can also be approximated using the MVF method:
σb0.2,lam = MV F + (1 − MV F)
Ex,pr
Em
α σb0.2,m (3.35)
σbult,lam = MV F + (1 − MV F)
Ex,pr
Em
α σbult,m (3.36)
where σb0.2,m is the metal layer bearing yield strength and σbult,m is the metal layer bearing
ultimate strength. The parameter α is used to correct the inﬂuence of geometrical parameters
on the bearing strength of the prepreg and metal layers.
These two equations were found to be suitable for the analysis and optimization of the bearing
properties of FMLs. A good value for the geometrical correction parameter α can usually be
obtained by curve-ﬁtting test data[10].

Chapter 4
Manufacturing of the Laminate
This chapter will discuss the manufacturing processes for certain existing FML’s as well as
the Carbon-fiber/Stainless Steel laminate panels that were used for this research. It will point
out some difficulties that were encountered during the production and how they were tackled.
4.1 Introduction
The manufacturing of Fiber Metal Laminates is often a rather complicated and time consum-
ing process. Every individual layer used in the FML must be prepared and often pre-treated
before it can be used in the final laminate. A very important factor in the manufacturing of
FML’s is the strength of the bond between the individual layers. Special care must thus be
taken to ensure adequate bonding between the constituents of the FML.
4.2 Manufacturing of Current Fiber Metal Laminates
This section will briefly go over the manufacturing processes of the four main types of FML
currently developed. These are GLARE (glass fibers and aluminium), ARALL (aramid fibers
and aluminium), CARE (carbon fibers and aluminium), and TiGr (carbon fibers and tita-
nium).
4.2.1 Manufacturing of GLARE
The manufacturing of GLARE begins with a surface treatment of the aluminium sheets. The
surface of the sheets is anodized and primed to provide a good bonding strength and good
protection against corrosion. For the anodization, either Chromic Acid Anodizing (CAA) or
Phosphoric Acid Anodizing (PAA) is used. To increase corrosion resistance, cladding can be
applied to the outer aluminium layers of the laminate.

18 Manufacturing of the Laminate
The laminate is cured in an autoclave at increased temperature and pressure. Depending on
the type of used matrix, this temperature is usually around 120◦C and the pressure is about
0.3 - 1.0MPa. During curing, residual stresses arise in the laminate due to different thermal
expansion coefficients of the aluminium and the prepreg. To counter these residual stresses,
the laminate is sometimes post-stretched after curing. This is mostly done for laminates
based on 7075-T6 and 7475-T76, which do not have as good fatigue properties as for example
2024-T3.
4.2.2 Manufacturing of ARALL
The manufacturing procedure of ARALL is essentially the same as the procedure for GLARE.
Post-stretching of the laminate is essential however, because of the low resistance to compres-
sive deformation of the aramid fibers.
4.2.3 Manufacturing of CARE
Galvanic corrosion
The manufacturing of CARE is also similar to the manufacturing of GLARE. The combi-
nation of carbon fibers with aluminium sheets however, presents the potential problem of
galvanic corrosion[4,5,31]. Despite the fact that the fibers are covered in an isolating epoxy
matrix, direct contact between the fibers and the aluminium can still occur, for example when
fasteners are used and holes are drilled in the material.
Three methods have been developed and tested so far to counter the occurrence of galvanic
corrosion:
• A primer is applied on the aluminium surface, immediately after the pre-treatment of
the aluminium sheets.
• The aluminium sheets are covered by a very thin (0.02 mm) thermoplastic layer based
on PolyEtherImide (PEI).
• A thin (0.1 mm) glass fiber prepreg layer is introduced to isolate the carbon fiber prepreg
layer on both sides.
The application of these isolating layers, although preventing galvanic corrosion, does reduce
the mechanical properties of the laminate. The fatigue properties of the isolated laminate are
still good however, when compared to GLARE or ARALL[5].
Care must be taken when making notches in the isolated material. It was found that in some
cases, the carbon and aluminium could make contact again in the edge area. Other potential
problem areas are the sheet edges or areas with scratches, fatigue damage or other incidental
damage[31].

4.3 Manufacturing of the Carbon-Fiber/Stainless Steel Laminate 19
Pre-treatment and curing
For the production of CARE, the aluminium sheets are given a pre-treatment consisting of
degreasing, pickling and CAA anodizing in order to obtain good adherence with the fiber
prepreg layers.
Curing of the laminate is also done in an autoclave. The curing cycle used for CARE depends
on the exact type of fiber and adhesive used, but in all cases involves an elevated pressure of
about 0.3 - 1.0 MPa (3 - 10 bar) and an elevated temperature of about 120 - 180 ◦C[5].
Post-stretching of CARE is generally not done due to the very low ultimate strain of the
carbon fibers[1].
4.2.4 Manufacturing of TiGr
The manufacturing of the TiGr laminate also begins with a pre-treatment of the titanium
sheets. Various methods can be chosen for this treatment. One of them is degreasing using an
alkaline solution, etching with either chromic or sulphuric acid and anodizing using NaTESi
anodizing[32]. Another method is degreasing followed by CAA anodizing or Sodium Hydroxide
Anodizing (SHA). Although SHA anodizing does result in a better durability than CAA
anodizing, the preferable anodizing method is CAA, because this pre-treatment provides a
higher bonding strength than the SHA method[33].
Curing of the laminates for the pre-treatment method as given by Medenblik[32] occurs in a
Fontijne hot plates press, at an elevated temperature of approximately 380 ◦C and a pressure
of 2.5 MPa. Curing for the method as given by Koos[33] is done in an autoclave at unspecified
temperature and pressure.
4.3 Manufacturing of the Carbon-Fiber/Stainless Steel Laminate
This section will describe how the Carbon-fiber/Stainless Steel laminate was manufactured,
what materials were used, what options for the pre-treatment of steel sheets were available
and which of those options was found to be optimal.
4.3.1 Choice of Materials and their Properties
For this project, the choice of materials was limited to what was readily available for use.
It turned out to be necessary to use a different type of steel and a different type of carbon
fiber/epoxy for the final bearing tests. However, since the initial bearing tests did not neces-
sarily need to be compared to the final bearing tests quantitatively, the difference in materials
was not an issue. The types of steel and carbon fiber/epoxy used are presented below.
Stainless steel
Due to the high specific weight of stainless steel, the best potential is found in thin sheets.
For the preliminary tests as well as the initial bearing tests, cold rolled Sandvik Nanoflex

stainless steel sheets (1kk91) were available. These sheets had a thickness of 0.08 mm. For
the final bearing tests, AISI 316L stainless steel sheets were used, with a thickness of 0.1 mm.
The properties of both types of steel can be found in Table 4.1.
Carbon fiber/epoxy
The carbon fiber/epoxy systems that were available for this project were HexPly
M21/35%/134/T700GC for the preliminary and initial bearing tests, and Delta-Preg M30SC
for the final bearing tests. Relevant data for these fiber/epoxy systems are given in Table 4.1
as well.
Material E11 E22 G12 ν12 X Y ρ
[GPa] [GPa] [GPa] [−] [MPa] [MPa] [g/cm3]
Nanoflex stainless steel[34] 185.0 185.0 69.6 0.33 1,400 1400 7.87
AISI 316L stainless steel[35] 193.0 193.0 0.33 560 560 8.00
HexPly M21/T700GC[36] 147 3.5 4.7 0.33 2,314 147 1.58
Delta-Preg M30SC[37] 175 7.8 - 0.33 2,990 51 1.53
Cytec FM94U adhesive film[38] NA NA NA NA NA NA 1.12
Table 4.1: Overview of material properties
4.3.2 Pre-treatment of the Steel Sheets
For good adhesion to the carbon fiber/epoxy layers, the steel sheets had to be pre-treated.
Several methods for the surface treatment of stainless steel sheets have been investigated. To
remove dirt, residue and other impurities the sheets can be degreased with organic solvents
or, more environment friendly, hydrous cleaners. Care must be taken not to leave any residue
on the sheets and not to re-apply any grease on already cleaned surfaces. In addition to the
above, treatment of the sheets in a bath with ultrasonic waves can be applied to improve
the degreasing process. Vapor degreasing is also mentioned as an effective method to remove
contaminations from the sheets, but this has not been tested in this research project.
Mechanical methods for further surface pre-treatment include brushing, grinding and blasting
of the sheets. These methods effectively increase the surface of the sheets, providing a larger
area for bonding with the adhesive. Brushing and grinding however can cause slight roughness
and could cause impurities to spread over the whole surface[4].
Chemical pre-treatments can also be used, but these are often costly and in addition, waste dis-
posal is becoming increasingly difficult. In addition, tests showed that chemical pre-treatments
do not result in the desired adhesion level as required for FML applications[4,40]. Furthermore,
in the case of thin steel sheets (0.1 mm), the level of etching required to improve adhesion
would be too detrimental to the sheets. A weight loss of 10% after 3 minutes of etching
has been registered. This pre-treatment method is therefore not really an option for these
sheets[40].
Coating of the sheet surface with primers after pre-treatment is also a possibility. This is
most useful for epoxy based adhesive films cured at elevated temperatures.

4.3 Manufacturing of the Carbon-Fiber/Stainless Steel Laminate 21
Several combinations of different surface treatments were investigated and it was concluded
that a degreasing process followed by sandblasting of the sheets resulted in the highest shear
strength of the bond[1]. A post-treatment with an 1% aqueous silane solution after sandblast-
ing has also proven to be effective[4,40].
For this research project, the stainless steel sheets were therefore first cleaned and degreased
using PFQD solvent/degreaser and PFSR cleaning agent. After cleaning and degreasing, the
steel sheets were sandblasted with corundum sand, sprayed from approximately 5 cm distance
from the sheets. This was done as evenly as possible, but because this is a manual process,
differences in the surface quality were hard to avoid. A Pneumix Pulsar III cabin was used
for the sandblasting process.
Due to the fact that the steel sheets are very thin, warping will occur if only one side of the
sheet is sandblasted. This makes lamination of the sheet impossible. Therefore it is essential
that the other side of the sheet is also sandblasted to compensate for the the warping. It is
advisable to attach the sheet in an open frame for this purpose, such that both sides can be
sandblasted easily without having to detach the sheet first. For this project, such a frame was
unavailable, hence the sheets were attached to a fixed plate using clamps. After sandblasting
one side, the sheets had to be detached, turned around, and re-attached to the plate. Due to
the warping, this was a time consuming process.
After sandblasting the sheets were degreased and cleaned once again before continuing with
the lay-up.
4.3.3 Lay-up of the Laminate Panels
After pre-treatment of the steel sheets, the carbon fiber/epoxy layers were prepared and
the laminate was made. An extra Cytec FM94U adhesive film layer was used between two
stainless steel sheets to provide bonding of the sheets. During the preliminary tests, this
adhesive layer was also used between the stainless steel and carbon fiber/epoxy layers of one
of the laminates in order to test if this would improve bonding. The properties of the adhesive
layer can be found in table Table 4.1 above.
In order to remove all air from between the laminate layers, the sheets were sealed in a
vacuum before being cured in the autoclave. However, despite the applied vacuum, pockets
of air could still remain in some places of the laminate. Since these air inclusions degrade the
quality of the laminate, it is necessary to test for the presence of these kind of defects after
curing of the laminate panels. This is usually done using an ultrasonic trough transmission
scan, also called a C-scan. This process is described in Section 4.3.5.
4.3.4 The Curing Cycle
Curing of the laminate panels was done in an autoclave. The laminate panels of the initial
bearing tests were kept at elevated temperature (120 ◦C) and pressure (0.7 MPa / 7 bar)
for a period of 8 hours. The temperature and pressure were increased linearly from ambient
conditions during a period of 50 minutes. Cooling down and dropping the pressure was also
done in a period of 50 minutes. For the panels of the final bearing tests, a different cycle was
used due to the different prepreg system used. The laminate was kept at elevated temperature

(120 ◦C) and pressure (0.6 MPa / 6 bar) for a period of 90 minutes, with linear warm-up and
cool-down periods of 45 minutes.
4.3.5 Testing for Inclusions and Contaminations
After the curing cycle, the laminate panels are tested for (air) inclusions and other con-
taminations using a C-scan machine. During this procedure, the panels are analyzed using
ultrasonic sound waves, which travel through laminar streams of water, passing the laminate.
Differences in thickness of the laminate or the presence of for example air pockets will cause
the sound waves to pass the laminate at different speeds. These differences in speed are reg-
istered, and defects can thus be visualized. Care can then be taken to avoid using these parts
of the panels for test specimens.
4.3.6 Production of the Test Specimens
After the sheets were fabricated, the individual test specimens had to be produced from the
panels. The edges of the panels were removed first, because the resin leaks a bit out of
the panel during curing, leaving sharp and thin edges which do not represent the laminate
structure well. Care was taken to maintain the correct alignment of the fibers in horizontal
and vertical direction during this process, because if the edges are cut off incorrectly, the fibers
will be oriented slightly off axis in the test specimens which eventually results in incorrect
test data. After removal of the edges, straight panels remained from which the final test
specimens were produced.
For the removal of the edges, as well as for the production of the final specimens, a Struers
Unitom 5 circular saw was used. The cutting blades used for this process were C54BF diamond
blades and the feed was set to 1 mm/s. The cuts produced with this method were clean and
straight, but it was necessary to remove any burrs that remained on the edges of the specimen
after sawing.
After the specimens had been fabricated to the proper dimensions, they were taken to the
workshop where holes for the bearing tests were drilled.
4.4 Concluding Remarks
Although care was taken to manufacture the laminate panels properly, small deviations in
procedure due to inexperience may have affected the quality of the final panels. Furthermore,
it is unfortunate that the choice of material combinations was limited to what was at hand.
If more materials would have been available, and time would not have been an issue, a more
detailed study should have been performed to determine what combination of steel and carbon
fiber prepreg would have been best.

Chapter 5
Testing the Laminate Concept
This chapter will discuss the tests that were performed on the Carbon-fiber/Stainless Steel
laminate in order to develop a better understanding of the bearing behavior of the laminate
concept.
5.1 Introduction
Presently, not a lot of research has been performed on Carbon-fiber/Stainless Steel laminates.
Dymáˇcek investigated the feasibility of the concept and found that the laminate performs well
in terms of mechanical properties, obtaining tensile strengths of 1560 MPa and more. With
respect to the bearing characteristics of the laminate, Dymáˇcek found that the lug bearing
strength is nearly proportional to the Metal Volume Fraction of the laminate. With larger
lug dimensions, he found that the prepreg layers contributed more to the bearing strength[1].
Shahinian also looked into the Carbon-fiber/Stainless Steel laminate concept and, focussing
on several different choices of fibers, investigated the bonding and mechanical properties of
the laminate. No tests were performed to determine the bearing properties of the laminate[7].
In order to further investigate the bearing characteristics of the Carbon-fiber/Stainless Steel
concept and in order to answer the questions posed in Chapter 1, a series of tests were
performed which will be described in this chapter. The test results and observed failure
mechanisms will also be presented.
5.2 Preliminary Tensile Tests
In order to determine which laminate configuration would be most suitable to continue this
research with, a series of preliminary tensile tests was performed on two different Carbon-
fiber/Stainless Steel laminate lay-ups. In addition to the test data, the production of these
panels provided valuable insight into the manufacturing process of the laminate as well.

24 Testing the Laminate Concept
5.2.1 Laminate Configuration
Two different panels were produced for the preliminary tests of the material. The configura-
tion of each of these panels is listed in Table 5.1.
Panel Configuration Layout Lay-up
P1A Cross Ply 2/1 [S/0/90/0/90]S
P1B Cross Ply 3/2 [S/0/90/45/45/90/0/S/0/90/45/45/90/0/S]
Table 5.1: Preliminary Tensile Test Laminate Configurations
Panel P1A consisted of eight carbon fiber layers which were sandwiched between two stainless
steel sheets. Panel P1B consisted of three steel sheets with six carbon fiber layers sandwiched
between each pair. Basic tensile tests were performed on specimens which were fabricated
from these laminates in order to determine their general behavior.
At this stage, the bonding quality of the steel and carbon was still uncertain. Therefore, an
extra layer of Cytec FM94U adhesive film[38] was introduced between the steel and carbon
layers in panel P1A to see if this would have an effect on the quality of the bond.
5.2.2 Specimen Geometry
The specimens created for the initial tests were to be tested for tensile strength only. Stan-
dards for tensile tests dictate that the specimens be created with a cut-out fillet area[39].
However, for the purpose of this preliminary test, rectangular specimens were deemed to be
sufficient. The omission of the fillet area greatly simplified the manufacturing procedure.
The specimens were specified to have a width of 10.0 mm and a length of 200 mm. Due to
slight inaccuracies during fabrication however, the final width of the manufactured specimens
varied between 10.4 and 10.6 mm.
5.2.3 Testing Conditions and Procedure
The specimens were tested on a Zwick/Roell 250kN static test machine under ambient con-
ditions. The specimens were tested with a loading rate of 6.0 mm/min until they failed.
5.2.4 Test Results
The results of the tests are presented in Appendix A in Figure A.1 and Figure A.2. A
summary of the results can be found in Table 5.2.
Panel E-modulus σult ult
[GPa] [MPa] [%]
P1A 53 900 1.8
P1B 70 720 1.2
Table 5.2: Tensile Test Results

5.3 Initial Bearing Tests 25
5.2.5 Brief Discussion of the Results
The preliminary test results showed that panel P1A performed about 20% better than panel
P1B in terms of ultimate tensile strength. This configuration also exhibited more consistency
in the results, with the load-displacement curves overlapping each other very closely. All of the
specimens of the P1A panel exhibited a bi-linear path until failure and the load-displacement
curves did not show any fluctuations. Failure of these specimens was instantaneous.
The load-displacement curves of the specimens of panel P1B were much more erratic, but
primarily after the point of initial damage. Three out of five specimens were able to withstand
additional loading after initial damage occurred in the specimen. The other two specimens
failed instantaneously, much like the specimens from panel P1A. In addition, the tests of
panel P1B had a much more explosive nature than the test of panel P1A. The specimens
were being torn apart and exhibited heavy buckling and delamination after failure.
This different behavior is most likely caused by the presence of fiber layers with a 45 degree
orientation.
The stiffness of panel P1B was found to be about 30% higher than the stiffness of panel P1A.
No significant differences between the bonding quality of the two panels (with or without the
extra layer of adhesive film) have been observed.
For the initial bearing tests it was decided to focus only on UD and CP configurations and
to not include any 45 degree layers.
5.3 Initial Bearing Tests
To determine if stainless steel strips could be used to improve the bearing strength character-
istics of the Carbon-fiber/Stainless Steel laminate, a series of bearing tests were performed
on the FML in both UD and CP configurations. Steel strips were added in three different
ways to try to determine an optimum combination of lay-up and reinforcement.
In this section the laminate configuration and the test setup of the initial bearing tests is
described. The results of these tests will be analyzed and discussed in Chapter 6.
The laminate panel configurations tested during the initial bearing tests are presented in
Table 5.3. The symbol S in the lay-up denotes the location of a stainless steel layer. The
symbol g denotes an extra layer of adhesive used to bond two layers of steel together. The
laminate density of all the panels was determined using the MVF approach.
The specimens created for these initial bearing tests were fabricated according to the ASTM
standard D953-87[39]. The specimens were specified to have a width of 48.0 mm and a length
of 150 mm. The hole for the bolt or pin was to have a diameter of 8.0 mm and was to be

Panel Configuration Lay-up Density
[g/cm3]
P2E-NR UD [S/0/0/0/0]S 2.16
P2B-NR CP [S/0/90/0/90]S 2.16
P2E-R2C UD [S/0/0/0/S/0]S 2.70
P2C-R2C CP [S/0/90/0/S/90]S 2.70
P2C-R2G CP [S/0/90/0/90/S/g/S/90/0/90/0/S] 2.57
P2B-R1M CP [S/0/90/0/90/S/90/0/90/0/S] 2.44
Table 5.3: Initial Bearing Test Laminate Configurations
positioned 24.0 mm off the edges. This would result in an E/D ratio of 3 and a W/D ratio
of 6.
However, due to the unavailability of correct fixtures for the specimens, the holes had to be
drilled to a diameter of 6.25 mm. Further, due to inaccuracies in the manufacturing process,
the width of the specimens varied between 48.4 and 48.5 mm. This resulted in a final E/D
ratio of approximately 3.8 and a W/D ratio of approximately 7.7. It is not believed that
these higher ratios negatively influenced the test results.
5.3.3 Test Conditions and Procedure
For the bearing tests a Zwick/Roell 250kN static test machine was used. The test procedure
followed was based on the ASTM standard D953-87[39]. The tests were carried out under
ambient atmospheric conditions with a loading rate of 1.3 mm/min.
Four specimens from each panel were tested under identical conditions. The specimens were
each pin loaded, but a finger-tight lateral constraint was provided to limit the occurrence of
buckling. This constraint was achieved by introducing aluminium filler plates between the
washers and the specimen.
Displacement of the actuator was allowed to continue until the load output began to plateau
or until a sharp drop in the load was observed.
5.3.4 Test Results
The load-displacement curves of the six bearing test series can be found in Appendix B in
Figure B.1 through Figure B.6. It can be observed that in most cases the load-displacement
curve follows a linear path up to a certain point, after which a slight drop in the load is
observed. The curve then continues to rise along an erratic line reaching a maximum value
after which the curve begins a zigzag pattern with sharp drops. Figure 5.1 shows a typical
bearing test curve where these features can be observed. The markers A through D below the
graph indicate the locations of the four stages of failure as will be described in Section 5.3.5.
Average values for the initial bearing tests are presented in Table 5.4 below, indicating three
Bearing Load levels:
• Damage Bearing Load

Figure 5.1: Typical bearing test curve
• Failure Bearing Load
• Maximum Bearing Load
The damage bearing load and failure bearing load were defined as suggested by Caprino[14]
and as presented in Section 2.4.2. The damage bearing load indicates the load level just
before the first drop in load is observed. The failure bearing load is the load level which is
reached thereafter, right before a sharp drop indicating severe damage in the specimen. The
definition of maximum bearing load was added to indicate the level of the maximum bearing
strength that was recorded during the test. This addition was necessary, since it was observed
that due to the occurrence of several sharp drops in the load, the failure bearing load was not
necessarily the maximum bearing load reached.
Some of the specimens suffered from slight delaminations or other visible defects. Whenever
these defects proved to be detrimental for the specimen test results, these results were not
included in the final results of the series.
Panel Damage B.L. St. Dev. Failure B.L. St. Dev. Max. B.L. St. Dev.
[N] [N] [N] [N] [N] [N]
P2E-NR 4,404 566 6,510 510 8,488 105
P2B-NR 4,788 471 6,233 403 9,323 722
P2E-R2C 6,359 376 10,528 255 12,335 1491
P2C-R2C 7,304 495 10,808 104 12,007 552
P2C-R2G 7,188 229 10,891 31 13,716 493
P2B-R1M 5,843 499 9,365 644 12,007 878
Table 5.4: Initial Bearing Test Results (Average Values)

5.3.5 Failure Modes
Observations
Most of the specimens tested have failed by a combination of bearing failure and a form of
shear-out. The shear-out was unexpected, since the E/D ratio that was chosen was thought
to be high enough for the specimens to fail by pure bearing failure. Any shear-out that
occurred however, was only present in the thin steel sheets. The composite layers did not
suﬀer from shear-out but failed by pure bearing failure. It can further be observed that the
growth direction of the steel shear-out is not simply vertical. Rather, it follows the path of the
buckling and turns slightly sideways in both directions. These two reasons seem to indicate
that the shear-out damage as observed is not the regular kind of shear-out as illustrated in
Figure 2.1.
Figure 5.2 illustrates four typical stages of failure as observed during the tests. These stages
can be described as follows:
• A. Damage onset - sheet buckling starts along the edge of the hole
• B. Damage growth - sheet buckling is clearly visible
• C. Local fracture - severe sheet buckling, partial shear-out
• D. Fracture growth - sheet buckling continues upwards, shear-out grows
Figure 5.1 shows at what point of the bearing tests these stages will approximately occur.
Figure 5.2: Typical failure stages during bearing failure (not the same specimen)
In several cases, heavy buckling of the sheet right above the hole caused delamination and
severe deformation above the buckled area, which in turn resulted in a crack forming at the
edge of the specimen. This process is illustrated with the series of pictures in Figure 5.3.
Examination of the cross-sections of the specimens revealed some information about the
behavior of the inner steel reinforcement sheet(s) when undergoing a bearing load. As can be

Figure 5.3: Crack formation at the top of the specimens (not the same specimen)
observed in Figure 5.4, the inner steel sheets seem to deform under the pressure of the buckling
that occurs at the outer steel sheets. The shape of the inner sheets is hence largely determined
by the locations of the folds in the outer sheets, which apply pressure onto the inner sheets
through the composite layers. This is the case for both single and double reinforcements, e.g.
one or two steel sheets in the center. In the case of double reinforcements, the two metal
layers practically behave like one layer (deforming symmetrically).
The occurrence of delamination and some matrix cracking can also be observed in the cross-
sections.
Figure 5.4: Close-up cross-section view of a reinforced specimen
Exceptions
Three exceptions were observed, in which the specimens failed with a diﬀerent failure mode
than the majority of the specimens. These exceptions are shown in Figure 5.5 and can be
described as follows:

• A. Straight shear-out of the metal sheet on one side of the sample.
• B. Crack in the metal sheet growing sideways instead of towards the edge.
• C. Shear-out cracks of the metal sheet stopping and turning sideways, with a crack
starting at the edge in the middle.
Figure 5.5: Failure mode exceptions
Relation of the failure mode to the laminate lay-up
After observation of the cross-sections of the specimens, it was noted that the presence of
the inner reinforcements can have an effect on the failure behavior of the specimens in some
cases. When sufficient load is applied, the inner sheets seem to buckle on their own close to
the edge of the specimen, with seemingly little or no influence from the outer steel layers.
However, this phenomenon was not observed in every test with inner reinforcements. It might
have been the result of poor bonding between the inner steel layers and the composite layers
around them, giving the inner steel layers more room to deform on their own under the applied
bearing load. The difference between this behavior and cases where the inner steel layers do
not buckle can clearly be observed in Figure 5.6. The figure depicts two specimens from the
same panel (thus with the same lay-up), both reaching a bearing load of approximately 12
kN.
5.4 Final Bearing Tests
After having gathered sufficient information from both the preliminary tensile tests as well as
the initial bearing tests, a series of final bearing tests were performed to further help answer
the questions posed in Chapter 1.

5.4 Final Bearing Tests 31
Figure 5.6: Close-up cross-section view showing buckling of the inner sheets
The laminate panel configurations produced for the final bearing tests are presented in Ta-
ble 5.5. The symbol S in the lay-up indicates the location of a stainless steel layer. In all
configurations, a layer of adhesive was used between two layers of steel to provide adequate
bonding. This is not explicitly denoted in the configurations below.
For the final bearing tests, only Cross Ply panels were used, with no variations in fiber
orientation. The outer fiber layers were always oriented in [0]-direction. The laminate density
of all the panels was determined using the MVF approach.
Panel Configuration Lay-up Lat. constraint Density
[g/cm3]
P3A-I CP [S/0/90/0/90/S]S no 2.87
P3A-II CP [S/0/90/0/S/90]S no 3.03
P3A-III CP [S/0/90/S/0/90]S no 3.03
P3A-IV CP [S/0/S/90/0/90]S no 3.03
P3A-V CP [S/S/0/90/0/90]S no 2.73
P3B CP [0/90/0/90/0/90]S yes and no 1.53
P3C-I CP [S/S/0/90/S/S/0/90]S yes 3.29
P3C-II CP [S/S/0/90/0/90/S/S]S yes 3.17
P3C-III CP [S/S/0/90/0/90/S]S yes 3.05
P3C-IV CP [S/0/90/S/S/0/90]S yes 3.20
P3D-I CP [S/0/90/0/90/S/0/90]S yes 2.62
P3D-II CP [S/0/90/0/90/S/0/90/0/90/S]S yes 2.65
P3D-III CP [S/0/90/0/90/S/0/90/0/90/S/0/90]S yes 2.53
P3E CP [S/0/90/0/90]S no 2.38
Table 5.5: Final Bearing Test Laminate Configurations
The two innermost steel reinforcement strips in panels P3C-I, P3C-II, P3C-IV, P3D-II and
P3D-III were longer than the outer strips in order to produce a more gradual increase of

thickness.
For the final bearing tests, a deviation was made from the ASTM standard as used for the
initial bearing tests. Since the available tooling limited the hole diameter to 6.25 mm, the
width of the specimens was decreased in order to approach the E/D ratio of 3. This resulted
in specimens with a width of 40.0 mm and a length of 150 mm. The hole was drilled at 20.0
mm distance from the edges. This resulted in an E/D ratio of 3.2 and a W/D ratio of 6.4.
5.4.3 Test Conditions and Procedure
For the final bearing tests the same equipment was used as for the initial bearing tests,
namely the Zwick/Roell 250 kN static test machine. The test procedure followed was also
based on the ASTM standard D953-87[39] and the tests were again carried out under ambient
atmospheric conditions with a loading rate of 1.3 mm/min.
After the initial bearing tests, it was observed that the most relevant point of the tests is the
point where the damage bearing load is reached (as defined in Section 5.3.4). The behavior
of the load-displacement curve after reaching this load is mostly erratic and not very relevant
from an engineering point of view, since most structures will usually be designed as to not
let the applied load surpass the damage bearing load. Therefore, for the final bearing tests,
emphasis was put on the analysis of this point.
Five specimens of each panel were tested under identical conditions for the final bearing tests.
All series of specimens were pin loaded as for the initial bearing tests, however this time some
of the series were tested without the lateral constraint as described in Section 5.3.3. Table 5.5
indicates which panels were tested with lateral constraint and which panels without. The goal
of this distinction was to observe what effect this would have on the damage bearing load and
on the failure mechanisms of the specimens occurring after the damage bearing load. The
composite-only panel (P2B) was tested both with and without lateral constraint.
For the tests without lateral constraints, displacement of the actuator was allowed to continue
until the first drop in the load was observed and the load output began to plateau afterwards.
The tests with lateral constraints were performed up to a displacement of about 5 mm or a
load of about 12-14 kN.
5.4.4 Test Results
Load-displacement curves of the fifteen final bearing test series can be observed in Appendix C
in Figure C.1 through Figure C.15.
The curves of the laterally restrained specimens show a similar behavior to the curves of the
initial bearing tests (Section 5.3.4). However, it can be observed that the zigzagging pattern
of the curves of the final bearing tests is less pronounced than those of the initial bearing tests.
Furthermore, the bearing load of the final bearing tests does not seem to flatten out as soon
as happened with the initial bearing tests. Instead, the load climbs to a higher point after

every drop. Therefore, and because the tests were often stopped before the load output began
to plateau, it is not useful to define the maximum bearing load level for the final bearing test
series with lateral constraint.
Figure 5.7: Typical bearing test curve (with lateral constraint)
The load-displacement curves of the test series without lateral constraint, however, display
a significantly different behavior from what has been observed before. Although the damage
bearing load level seems to be as high as for the tests with lateral constraint, the load drop
after the damage bearing load level is much more severe. Furthermore, after this drop, the
load level seems to be unable to recover to a point higher than the damage bearing load in
most of the cases. Instead, the load level either continues to drop gently, plateaus, or climbs
slowly until reaching another drop. This behavior seems to depend on the configuration of
the laminates however, and will be discussed in more detail in Chapter 6.
Typical load-displacement curves displaying the features as discussed above for both later-
ally constrained and unconstrained specimens can be found in Figure 5.7 and Figure 5.8
respectively.
Average values of the damage bearing load and failure bearing load (when applicable) are
presented in Table 5.6.
5.4.5 Failure Modes
Observations
The failure mechanisms observed during the final bearing tests were slightly different from the
failure mechanisms that were observed during the initial bearing tests. Most of the specimens
did fail by bearing failure, but a form of shear-out of the sheets was not observed. Instead
however, a small crack could be observed which started to grow at the edges of the hole of
some specimens (mostly the laterally constrained specimens), but this crack never reached
large proportions.

Figure 5.8: Typical bearing test curve (without lateral constraint)
Buckling that occurred during the final bearing tests was also different from the buckling
observed during the initial bearing tests. The steel layers buckled more smoothly and in
larger waves, in contrast to the compact and small buckling waves that were observed during
the initial bearing tests. Furthermore, the buckling spread out more towards the sides of the
specimen, whereas during the initial bearing tests buckling was effectively only present in the
steel located between the shear-out cracks.
A possible reason for the lack of shear-out of the steel is that for the final bearing tests a
different and slightly thicker steel type was used (Section 4.3.1), giving the material more
resistance. The different material properties most likely affected the buckling behavior as
well.
Differences between the laterally constrained and unconstrained specimens also exist. When
looking at the unconstrained specimens, very little buckling above the hole can be observed.
Instead, the outer steel layers fold outwards due to the lack of any constraint. Four photo’s
illustrating this typical behavior for the unconstrained specimens are presented in Figure 5.9.
The constrained specimens showed more buckling above the hole and, as mentioned above,
a small crack could often be observed which started to grow at the edges of the hole. In
some cases, evidence of shear-out of the inner laminate layers could be observed. These
typical features of the failure modes of the laterally constrained specimens can be observed
in Figure 5.10.
For the composite-only panel (P3B), tests were carried out both with and without lateral
constraints. Although the specimens with lateral constraints did have better performance,
the failure mechanisms observed were similar for both cases. The specimens without lateral
constraints do not seem to exhibit a tendency to deform out of plane, like the panels with
steel sheets did. Both the constrained and unconstrained specimens suffered from shear-out
in some cases as well. Figure 5.11 illustrates typical failure modes for the composite panel.
Regarding tests that were allowed to continue for a longer period of time (larger displacement),

Panel Damage B.L. St. Dev.
[N] [N]
P3A-I 4,417 156
P3A-II 4,215 74
P3A-III 4,043 162
P3A-IV 3,929 157
P3A-V 3,821 127
P3B 4,333 181
P3B (constrained) 4,507 172
P3C-I 6,428 418
P3C-II 6,045 282
P3C-III 5,355 216
P3C-IV 5,556 157
P3D-I 5,933 218
P3D-II 9,275 200
P3D-III 11,725 146
P3E 3,156 302
Table 5.6: Final Bearing Test Results (Average Values)
it can be said that the same failure mechanisms occurred, but with amplified features. For
the laterally constrained cases, buckling was more severe and reached all the way up to the
edge of the specimen. Furthermore, the cracks at the edges of the hole were larger, but they
still did not result in shear-out-like features. For the laterally unconstrained cases, damage
was understandably more severe as well, with cracks at the edge of the hole present in some
cases. This behavior can be observed in Figure 5.12.
Considering the cross-sections of the specimens of the final bearing tests, it can be observed
that the behavior of the steel layers is different from what was observed earlier at the initial
bearing tests. The outer metal layers of the laminate appear to have a strong tendency to
bend outwards near the hole. It does not seem to matter whether the specimens were laterally
constrained or not, as all of the specimens that were tested seem to exhibit this behavior.
Figure 5.9: Typical observed failure modes (unconstrained specimens)

Figure 5.10: Typical observed failure modes (laterally constrained specimens)
Figure 5.11: Typical observed failure modes (composite-only panel)
Figure 5.13 provides a good illustration of this characteristic.
It can further be noted that the eﬀect of the bearing damage does not seem to extend as far
towards the edge of the specimens as was the case for the initial bearing tests. The steel layers
buckle mostly near the hole, leaving the material closer to the edge intact. For specimens of
the initial bearing tests, it could be observed that although the steel buckled less excessively,
the damage propagated much further towards the edge of the laminate.
A more detailed analysis of the cross-sectional behavior of the specimens is presented in
Chapter 6.
Exceptions
No noteworthy exceptions were observed during the ﬁnal bearing tests.

Figure 5.12: Specimen damage after large displacements1
Figure 5.13: Cross-section views showing outward bending of steel layers
1
P3C-III-B3: 6.8 mm laterally constrained
P3A-V-B1: 5.3 mm unconstrained
P3E-B1: 7.2 mm unconstrained

Chapter 6
Discussion of the Test Results
This chapter will discuss the results of initial and final bearing tests as described in Chapter 5.
The test results and observations will be analyzed in order to find answers for the questions
posed in Section 1.3. Furthermore, the results of the tests will be compared to the findings
of Chapter 2.
6.1 Effect of the Laminate Configuration
The results of the initial bearing tests confirmed that laminate panels with a cross ply config-
uration (P2B-NR) performed better than identical panels with a unidirectional configuration
(P2E-NR), as was also observed in the literature, see Section 2.3.2. The average damage
bearing load level for the CP panel was found to be about 9% higher than for the UD panel.
Qualitatively, it can be added that panels with a CP configuration showed more consistency
in the results, meaning less scatter was observed in the values for individual specimens of the
same series. This makes the behavior of CP panels more predictable.
For the final bearing tests, only CP panels were used.
6.2 Effect of the Addition of Steel
To determine if the addition of steel reinforcement strips to the laminate has a positive effect
on the damage bearing strength, the test results of the reinforced and non-reinforced panels
have to be compared. To ensure a fair comparison, with the effect of laminate thickness and
laminate density canceled out, the test results have to be normalized first by dividing the
Damage Bearing Load (DBL) by both the laminate thickness and density. The following
equation illustrates this process:
DBLnormalized =
DBLabsolute
ρlam · tlam
(6.1)

40 Discussion of the Test Results
After the normalized damage bearing loads have been calculated, the values are related to
each other. This results in what will be referred to as the relative damage bearing strength.
Figure 6.1 shows the effect of the addition of extra steel reinforcement strips to the center
of the laminate on the relative damage bearing strength. The values of the damage bearing
load are related to the results of the equivalent, non-reinforced versions of the panels which
had only steel on the outside.
Figure 6.1: The effect of the addition of steel strips on the damage bearing strength
It can be observed that the addition of steel reinforcement strips to the center of the laminate
improved the damage bearing strength of the panels during the initial bearing tests, but
negatively influenced the damage bearing strength of the panels during the final bearing tests.
The improvement was about 2% for panel P2B (CP configuration, one steel strip added) and
about 4% for panel P2E (UD configuration, two steel strips added). Panel P3A-I, with a CP
configuration and two steel strips added, performed about 6% worse than panel P3E.
This result can be attributed to the fact that the materials that were used for the final bearing
tests differed significantly from the materials that were used for the initial bearing tests. Refer
to Table 4.1 for an overview of the material properties. The steel used for the final bearing
tests (AISI 316L) had a slightly higher density and was 25% thicker than the steel used for
the initial bearing tests (Nanoflex). Moreover, the strength of the AISI 316L steel is about
60% lower than the strength of the Nanoflex steel. In addition, the strength of the carbon
fiber prepreg used for the final bearing tests (Delta-Preg M30SC) was about 29% higher in
X-direction than the strength of the prepreg used during the initial bearing tests (HexPly
M21). The ratio the Young modulus of the steel divided by the modulus of the prepreg was
similar for both the initial (1.25) and final (1.10) bearing tests.
The effect of the higher steel density is negligible, but the larger thickness of the AISI 316L
steel can be held accountable for a lower relative damage bearing strength of about 3%. The
fact that the AISI 316L steel is much weaker than the Nanoflex steel and that the Delta-Preg

6.2 Effect of the Addition of Steel 41
M30SC prepreg is stronger than the HexPly M21 prepreg, should account for the remainder
of the apparent poor performance of added steel to the laminate during the final bearing tests.
The contrast between the strength of both materials of the laminate was simply too large,
hence the addition of steel reinforcement strips did not seem to benefit the damage bearing
load in case of the final bearing tests.
Figure 6.2: Relative damage bearing load of all the panels of the final bearing tests
A comparison of the relative damage bearing strength of the panels of the final bearing tests
with the composite-only panel also supports this result. Figure 6.2 clearly indicates that the
performance of the composite panel was superior in the case of relatively thin panels. (The
effect of the thickness of the laminate on the damage bearing load is analyzed in Section 6.4).
For the initial bearing tests, unfortunately no composite-only panels were tested.
When the relative damage bearing load values for both test series are plotted against the
MVF values of the laminate panels, as is done in Figure 6.3, the positive effect of steel for
the initial bearing tests can clearly be observed in contrast to the negative effect of steel for
the final bearing tests. For this plot, the results of the initial bearing tests were related to
the results of the composite-only panel of the final bearing test series.
Figure 6.3: Metal volume fraction plotted against the relative damage bearing load
It should be noted at this point, that when the MVF is plotted against the absolute damage

42 Discussion of the Test Results
bearing strength, the data of the final bearing tests becomes too scattered to indicate a clear
trend. The data for the initial bearing tests does show a trend, but more tests would have to
be performed to determine if this trend fits with damage bearing strength values of composite
only laminates (MVF = 0) or specimens made from just steel (MVF = 1).
6.3 Effect of the Distribution of the Steel Strips
To determine the effect of the distribution of the steel strips within the panels on the bearing
strength, the test results of various panels can be compared:
• P2C-R2G and P2C-R2C can be compared, because they have the same steel/carbon
content but differently located inner steel sheets.
• P3A-I through P3A-V were specifically constructed with a varying location of the inner
steel sheets and are thus excellent for comparison.
• P3C-I through P3C-IV were constructed like panels P3A-I and P3A-III, but feature
double metal sheets. These panels will also be compared in this discussion.
The comparison of the P2C and P3A panels is presented in Figure 6.4, with the bars indicating
the absolute damage bearing load. Below the graph, the distribution of the steel sheets within
the laminate is illustrated.
Figure 6.4: The effect of the distribution of the steel strips on the damage bearing strength
From this comparison it is clear that there is an effect of the distribution of the steel strips on
the damage bearing load. The optimal distribution seems to be with the steel strips placed in
the center of the stack. When the strips are moved away from the center, the average damage
bearing load decreases linearly by about 2 to 4% per step.

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