i
ANCHORAGE SYSTEMS IN CONCRETE
STRUCTURES STRENGTHENED WITH
CARBON FIBER REINFORCED POLYMER
COMPOSITES
By
ROBIN KALFAT
A thesis submitted in fulfilment of the requirements for the degree of
Doctor of Philosophy
DEPARTMENT OF ENGINEERING AND INDUSTRIAL SCIENCES
SWINBURNE UNIVERSITY OF TECHNOLOGY
February 2014
Abstract
ii
ABSTRACT
Over the last two decades, extensive research has demonstrated the effectiveness of
externally bonded (EB) fiber-reinforced polymer (FRP) composites for strengthening
and repairing reinforced concrete (RC) structures. The main advantages of EB FRP for
strengthening applications, when compared to strengthening using traditional
engineering materials such as steel, are their high strength-to-weight ratio (up to ten
times stronger than steel and about 20% of the weight) and high corrosion resistance.
A commonly documented failure mode of FRP strengthened RC is premature
debonding, which generally occurs at fiber elongations well below the tensile strength
of the FRP. Failure by debonding is usually rapid and represents a significant
underutilisation of the materials strength properties. Design guidelines around the world
are strongly influenced by such behaviours and adopt a preventative approach by
limiting the design strain in the FRP to a level where debonding will not occur. A
logical means to improve the performance of externally bonded FRP by preventing end
debond is by anchorage. However, research in the field of anchorage systems has been
very limited to date. The demand to improve the efficiency of FRP systems, together
with the shortcomings in available research has inspired the present dissertation, which
will consist of experimental and numerical studies to develop a novel anchorage system
to address the present shortcomings of premature FRP debonding.
A state of the art literature review was conducted on available research in the area of
FRP anchorage systems. This provided a comprehensive summary which resulted in a
database where anchorage effectiveness factors were assigned to each anchor type as
part of an assessment of anchor performance. Of the many anchorages presented and
discussed, metallic anchorages were demonstrated to be the most effective form of
anchorage when using the maximum fiber elongation prior to failure as the sole
evaluation criterion. This was followed by non-metallic anchors such as: U-jackets and
spike anchors for use in flexural strengthening. Of the anchors used for shear
strengthening, flange embedment and the use of FRP spike anchors proved the most
efficient.
Abstract
iii
The present study commenced with an attempt to improve the anchorage strength by an
improvement of the substrate properties to which the FRP is bonded. Preliminary
investigations demonstrated that the strength of the concrete substrate is a key factor
affecting the delamination mode and overall bond strength. The introduction of a
mechanical chase cut into the concrete over the anchorage length was demonstrated as
an effective method to improve the strength of the concrete substrate, resulting in higher
FRP elongations, bond stresses, slips and load carrying capacities. The effect of the
chase was a 95-100% increase in joint capacity, 118% increase in bond stress and 83-
93% increase in the maximum strain level reached prior to failure.
Although improving the substrate properties showed promising results, non-destructive
anchorages were devised, resulting in a further experimental study using unidirectional
and (±45º) bidirectional fabric patch anchors. The anchorages tested were successful in
improving the FRP strain utilisation by up to 195%. The use of (±45º) bidirectional
fabric patch anchors, applied to the ends of FRP laminates resulted in a more efficient
distribution of FRP-adhesive stresses over a greater area of concrete. The remainder of
the experimental work (stage 2) focused on further developing the concept (±45º)
bidirectional fiber patch anchors. The influence of patch anchor size was investigated,
together with laminate thickness and width. The study concluded that patch anchor
lengths of 250mm or less, exhibited slippage at a lower load and that lengths of 300mm
were preferred in order to fully engage the anchor. By examining the strain distributions
within the bidirectional fibers it was found that laminates could be spaced as closely as
250mm without any reduction of anchorage strength.
Numerical finite element simulations were conducted which were able to capture the pre
peak and post peak response of the patch anchored joints to a high level of accuracy,
once calibrated with the numerical data. Parametric studies on concrete strength were
performed to expand the experimental data, resulting in an approximately linear
relationship between the concrete compressive strength and the maximum laminate
strain achieved prior to debond. Both the experimental data from stages 1 and 2, as well
as the information derived from the finite element simulations were used to develop a
theoretic anchorage strength model for the (±45º) bidirectional fiber patch anchored
joints. The model was capable of offering anchorage strength predictions for alternative
Abstract
iv
material and geometrical properties and was verified with the existing experimental and
numerical data.
Acknowledgements
v
ACKNOWLEDGMENTS
I would like to thank my supervisors, Prof Riadh Al-Mahaidi and Prof. John Wilson for
their guidance and support throughout this extensive project. May we continue to share
future collaborations and strong working friendships.
I would further like to acknowledge the Westgate Bridge Strengthening alliance for
their financial contributions and support for this research. For the assistance in the
preparation and construction of (stage 1) experimental specimens, I would like to thank
Dr Matthew Sentry and gratefully acknowledge the services provided by the
Department of Civil Engineering at Monash University and the laboratory staff
members: Alan Taylor, Kevin Nievaart and Long Goh.
I wish to thank the staff at the Civil Engineering SMART Structures Laboratory at
Swinburne University. The completion of experimental stage 2 of the experimental
works would not have been possible without the hard work and assistance of Michael
Culton, Kia Rasekhi and Sanjeet Chandra. For contributions and assistance with
overcoming various obstacles throughout this experimental component, I would like to
acknowledge Senior Test Engineer Graeme Burnett.
To the staff and postgraduate community in the faculty of Engineering and Industrial
sciences, thank you for your great friendships which made my PhD more than a time of
study. Finally, my greatest appreciation is reserved for my family, whose support and
encouragement is a source of strength and inspiration. This dissertation is a testament of
their support.
Declaration
vi
DECLARATION
I hereby declare that this thesis contains no material accepted for any other
degree or diploma in any university. To the best of my knowledge, this thesis contains
no material previously written or published by another person, except where due
acknowledgment is made in the text.
ROBIN KALFAT
Table of Contents
vii
TABLE OF CONTENTS
ABSTRACT……………………………………………………………………………………………………………….….ii
ACKNOWLEDGEMENTS………………………………………………………..………………………….……...v
DECLARATION………………………………………………………………………………..………………..………..vi
TABLE OF CONTENTS……………………………………………………………………………………….....…..vii
LIST OF FIGURES…………………………………………………………………………………………..………..…xii
LIST OF TABLES……………………………………………………………………………………………………..…xx
NOTATION AND ABBREVIATION…………………………………………….………………… ….…xxii
PART I CHAPTER 1 – INTRODUCTION ................................................................................... 1
1.1 General background............................................................................................ 1
1.2 Research problem and aims ................................................................................ 2
1.3 Thesis outline ..................................................................................................... 4
CHAPTER 2 - LITERATURE REVIEW OF FRP ANCHORAGE SYSTEMS IN
CONCRETE INFRASTRUCTURE ................................................................................. 5
2.1 Introduction ........................................................................................................ 5
2.2 Mechanisms of FRP failure and debonding for flexurally strengthened
members ........................................................................................................................ 7
2.3 Anchorage devices for FRP reinforcement used to strengthen members in
flexure ............................................................................................................................ 8
2.3.1 FRP U-jacket anchors ................................................................................. 8
2.3.2 Inclined U-jacket orientations ................................................................... 11
2.3.3 Prestressed U-jackets ................................................................................ 13
2.3.4 Metallic Anchorage Systems .................................................................... 14
2.3.5 FRP Anchors ............................................................................................. 18
2.3.6 Evaluation of FRP anchors used to strengthen members in flexure. ........ 21
2.4 Mechanisms of FRP failure in shear retrofit applications ................................ 26
2.5 Anchorage devices for FRP reinforcement used to strengthen members in
shear .......................................................................................................................... 26
Table of Contents
viii
2.5.1 Mechanically fastened metallic anchors in shear and torsion applications ..
................................................................................................................... 27
2.5.2 Anchorage of FRP through concrete embedment ..................................... 31
2.5.3 FRP spike anchors in shear applications ................................................... 32
2.5.4 Evaluation of FRP anchors used to strengthen members in Shear ........... 35
2.6 Conclusions ...................................................................................................... 39
CHAPTER 3 – LITERATURE REIEW OF FRP-TO-CONCRETE BOND
BEHAVIOUR ................................................................................................................. 41
3.1 Test set-ups and failure modes ......................................................................... 41
3.2 Bond transfer mechanism ................................................................................. 43
3.2.1 Parameters influencing bond strength ....................................................... 44
3.3 Modelling FRP Debonding .............................................................................. 46
3.3.1 Bond slip models ....................................................................................... 46
3.3.2 Concrete fracture energy methods ............................................................ 50
3.3.3 Bond strength models................................................................................ 52
CHAPTER 4 - EXPERIMENTAL INVESTIGATION INTO FRP ANCHORAGE
SYSTEMS UTILISING A MECHANICALLY STRENGHTNED SUBSTRATE ....... 55
4.1 Introduction ...................................................................................................... 55
4.2 Specimen Design .............................................................................................. 56
4.2.1 The Mechanically Strengthened Substrate Anchor ................................... 56
4.3 Test Preparation and Material properties ......................................................... 57
4.3.1 Control Specimen ...................................................................................... 58
4.3.2 Anchor Type 1 .......................................................................................... 58
4.3.3 Experimental Setup ................................................................................... 62
4.3.4 Instrumentation and loading procedure..................................................... 64
4.4 Experimental Results ........................................................................................ 64
4.4.1 Quality control tests .................................................................................. 64
4.4.2 Failure modes ............................................................................................ 66
4.4.3 Tilt ............................................................................................................. 68
4.4.4 FRP strain distributions ............................................................................. 69
4.4.5 Experimental bond slip curves .................................................................. 72
4.4.6 Effective strain in FRP laminates used in design ...................................... 75
4.5 Summary .......................................................................................................... 76
Table of Contents
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CHAPTER 5 - EXPERIMENTAL INVESTIGATION INTO FRP ANCHORAGE
SYSTEMS UTILISING UNIDIRECTIONAL AND BI-BIRECTIONAL FIBER
PATCH ANCHORS ....................................................................................................... 78
5.1 Introduction ...................................................................................................... 78
5.2 Specimen Design .............................................................................................. 78
5.3 Test Preparation and Material properties ......................................................... 80
5.3.1 Anchor Type 2: ......................................................................................... 81
5.3.2 Anchor Type 3: ......................................................................................... 82
5.3.3 Anchor Type 4: ......................................................................................... 84
5.3.4 Anchor Type 5: ......................................................................................... 85
5.3.5 Anchor Type 6: ......................................................................................... 87
5.4 Experimental Results ........................................................................................ 88
5.4.1 Failure modes ............................................................................................ 88
5.4.2 FRP strain distributions along length of laminate ..................................... 94
5.4.3 Load – Displacement curves ..................................................................... 98
5.4.4 Experimental bond slip curves ................................................................ 103
5.4.5 Strain in bidirectional fibers .................................................................... 107
5.5 Summary ........................................................................................................ 109
CHAPTER 6 – EXPERIMENTAL INVESTIGATION INTO THE SIZE EFFECT OF
BIDIRECTIONAL FIBER PATCH ANCHORS ......................................................... 111
6.1 Introduction .................................................................................................... 111
6.2 Experimental Program .................................................................................... 111
6.2.1 Specimen Design ..................................................................................... 111
6.2.2 Specimen preparation .............................................................................. 116
6.2.3 Experimental Setup ................................................................................. 117
6.2.4 Test Preparation and Material properties ................................................ 117
6.2.5 Instrumentation and loading procedure ................................................... 118
6.2.6 Image correlation photogrammetry ......................................................... 118
6.3 Experimental Results ...................................................................................... 121
6.3.1 Quality Control Tests .............................................................................. 121
6.3.2 Failure Modes ......................................................................................... 125
6.3.3 Overview ................................................................................................. 127
6.3.4 Load Deformation curves ........................................................................ 130
Table of Contents
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6.3.5 FRP strain distributions along length of laminate ................................... 134
6.3.6 Strain in Bidirectional fibers ................................................................... 138
6.3.7 Experimental Bond Slip Curves .............................................................. 143
6.4 Conclusion ...................................................................................................... 145
CHAPTER 7 – FINITE ELEMENT MODELLING OF UNIDIRECTIONAL AND
BIDIRECTIONAL FIBER PATCH ANCHORS ......................................................... 147
7.1 Introduction .................................................................................................... 147
7.2 The Proposed Finite Element Model .............................................................. 147
7.2.1 Modelling of concrete ............................................................................. 148
7.2.2 Modelling FRP Patch Anchors ............................................................... 151
7.2.3 Modelling steel reinforcement ................................................................ 152
7.2.4 Modelling FRP-to-Concrete Interface .................................................... 152
7.2.5 Solution strategies ................................................................................... 155
7.2.6 Element type for the concrete prism ....................................................... 155
7.3 Boundary Conditions ...................................................................................... 157
7.4 Numerical and Experimental Results ............................................................. 157
7.4.1 Type 0 – Control specimen Results ........................................................ 158
7.4.2 Type 5 – Bidirectional fabric specimen Results ..................................... 160
7.4.3 Type 2 – Unidirectional fabric specimen Results ................................... 165
7.5 Parametric studies........................................................................................... 168
7.5.1 Sensitivity to mesh size ........................................................................... 168
7.5.2 Sensitivity to fracture energy .................................................................. 171
7.5.3 Sensitivity to adhesive stiffness .............................................................. 174
7.5.4 Sensitivity to concrete strength ............................................................... 176
7.6 Conclusion ...................................................................................................... 179
CHAPTER 8 – FINITE ELEMENT INVESTIGATION INTO THE SIZE EFFECT OF
BIDIRECTIONAL FIBER PATCH ANCHORS ......................................................... 180
8.1 Introduction .................................................................................................... 180
8.2 The Proposed Finite Element Model .............................................................. 180
8.2.1 Modeling of concrete .............................................................................. 181
8.2.2 Modeling FRP Patch Anchors................................................................. 182
8.2.3 Modeling steel reinforcement ................................................................. 182
8.2.4 Modeling FRP-to-Concrete Interface...................................................... 182
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8.3 Results of non-linear finite element analyses ................................................. 184
8.3.1 Crack patterns and failure modes ............................................................ 184
8.3.2 Peak loads ............................................................................................... 191
8.3.3 FRP strain distributions along length of laminate ................................... 192
8.3.4 Strain in Bidirectional fibers ................................................................... 199
8.4 Parametric studies ........................................................................................... 200
8.4.1 Sensitivity to concrete strength ............................................................... 200
8.5 Summary ........................................................................................................ 203
CHAPTER 9 – DEVELOPMENT OF PATCH ANCHOR PREDICTION MODEL .. 204
9.1 Introduction .................................................................................................... 204
9.2 Assessment of prediction models ................................................................... 204
9.3 Parameters influencing an anchorage prediction model ................................. 205
9.3.1 Concrete Strength .................................................................................... 205
9.3.2 FRP width ............................................................................................... 207
9.3.3 FRP spacing ............................................................................................ 208
9.3.4 FRP thickness .......................................................................................... 209
9.3.5 Anchorage length .................................................................................... 210
9.4 Proposed anchorage strength model ............................................................... 210
9.5 Verification of the proposed model ................................................................ 211
9.6 Summary ........................................................................................................ 214
CHAPTER 10 – CONCLUSION .................................................................................. 215
REFERENCES .............................................................................................................. 218
LIST OF PUBLICATIONS…................................………………………………………………………………...229
List of Figures
xii
LIST OF FIGURES
Figure 2.1 - Types of CFRP debonding (adapted from Pham and Al-Mahaidi 2004) ...... 7
Figure 2.2 – U-shape anchoring method 45 degree ........................................................ 11
Figure 2.3 – Prestressing system for FRP ligatures (modified from Pham and Al-
Mahaidi 2006) ................................................................................................................. 14
Figure 2.4 - (a) Typical FRP plate anchored using permanent mechanical anchorage
device (b) schematic of typical test setup ...................................................................... 15
Figure 2.5 - (a) mechanical fastener; (b) predrilled holes; (c) Details of the HB-FRP
system; adapted from (Wu and Huang 2008). ................................................................ 17
Figure 2.6 - (a, b, c) Anchor construction and installation of FRP anchors (reprinted
from Engineering Structures, Vol. 33, No. 4, Smith, ST, Hu, S, Kim, SJ & Seracino, R
2011, “FRP-strengthened Rc slabs anchored with FRP anchors”, Pages 1075–1087,
April 2011, with permission from Elsevier); (d) test setup (single lap) (reprinted from
Construction and Building Materials, FRPRCS9 Special Edition, H.W. Zhang, S.T.
Smith, S.J. Kim, “Optimisation of carbon and glass FRP anchor design”, Pages 1–12,
June 2012, with permission from Elsevier); (e) generic load-slip response of FRP-to-
concrete joint anchored with bow-tie anchor; (f) joint strength enhancement (above
unanchored control) [modified from Zhang and Smith (2012b)] ................................... 20
Figure 2.7 - Implemented strengthening schemes (a) U-jacket; (b) Extended U-jacket;
adapted from (Deifalla and Ghobarah 2010) .................................................................. 28
Figure 2.8 - View Anchorage System with discontinuous steel anchorages, adapted from
(Ortega et al. 2009). ........................................................................................................ 30
Figure 2.9 – Steel anchorage schemes for strengthening of T-beams in shear; adapted
from (Aridome et al. 1998). ............................................................................................ 31
Figure 2.10 - (a) Typical FRP plate embedded 150mm into beam side with epoxy resin
(b) Typical schematic of typical test setup ...................................................................... 32
Figure 2.11 - Typical details of FRP spike anchors applied to shear applications ......... 33
Figure 3.1 - Different set-ups for shear-lap tests: a) Double pull-pull test; b) Single .... 42
Figure 3.2 – FRP-to-concrete joint typical bond stress distribution (a) top view (b) strain
distribution along FRP and (c) shear stress distribution along FRP (Lee, 2003)............ 43
Figure 3.3 – Typical bond slip curve .............................................................................. 48
Figure 4.1 - Control specimen geometry (WG9) configuration of strain gauges; .......... 58
List of Figures
xiii
Figure 4.2 - Anchorage type 1 specimen geometry (WG1 & WG2) (a) configuration of
strain gauges; (b) chase details and installation of N24 reinforcement bar (c) section
through chase. ................................................................................................................. 61
Figure 4.3 - Construction process of Type 1 Anchorage Specimen; (a) surface of
concrete block coated with MBRACE primer and centralisers for N20 reinforcement bar
located within chase; (b) profiling of laminate adhesive (as per manufacturers
specification) to the surface concrete block over reinforcement bar; (c) specimen curing
at an elevated temperature of 41°C. ................................................................................ 61
Figure 4.4 - Specimen testing rig details (a) configuration of test rig (front view); (b)
configuration of test rig (side view) ................................................................................ 63
Figure 4.5 - Specimen testing rig clamped to Baldwin testing machine (a) configuration
of test rig (front view); (b) configuration of test rig (rear view) ..................................... 63
Figure 4.6 - Adhesion testing and pressure gauge reading from test (TYFO BCC ±45°
fabric) showing failure within concrete. ......................................................................... 66
Figure 4.7 - Failed Control Sample (WGB9) (a) complete debonding of laminate from
concrete surface; (b) concrete surface post debonding of laminate (c) de bonded
laminate strip; (d) real time load, strain and ARAMIS photogrammetry recordings
during testing phase......................................................................................................... 67
Figure 4.8 - Testing of WGB1 (a) specimen ready for testing; (b) concrete rupture at
adhesive concrete interface; (c) debonded laminate strip. .............................................. 68
Figure 4.9 - Strain vs distance along Laminate; (a) Control specimen (WG9); (b) Type 1
- Anchorage specimen (WG1) ; (c) Type 2 - Anchorage specimen (WG2) ................... 69
Figure 4.10 - Load vs strain distribution, control specimen (WG9); .............................. 70
Figure 4.11 - Load vs strain distribution (a) Type 1 - Anchorage specimen (WG1); (b)
Type 1 - Anchorage specimen (WG2) ............................................................................ 71
Figure 4.12 - Bond-slip curves (a) Control specimen (WG9) with fitted curve following
Popovics equation; (b) Type 1 - Anchorage specimen (WG1) ....................................... 74
Figure 5.1 - Anchorage types 2 -5 applied to a box girder bridge. ................................. 79
Figure 5.2 - Anchorage types 2 specimen geometry and material properties ................. 81
Figure 5.3 - Construction process of Type 2 Anchorage Specimen; (a) Placement and
rolling out of voids of the first layer of MBRACE CF140, positioned 90° to the
direction of the laminate strip; (b) Profiling of laminate adhesive (as per manufacturers
specification) to the surface concrete block over MBRACE CF140 fabric and
List of Figures
xiv
application of application of MBRACE saturant; (c) Placement of second (top) layer of
MBRACE CF140 sheet to concrete block directly over location of first layer. ............. 82
Figure 5.4 - Anchorage types 2 and 3 applied to a box girder bridge. ............................ 83
Figure 5.5 - Anchorage types 3 specimen geometry and material properties. ................ 83
Figure 5.6 - Construction process of Type 3 Anchorage Specimen; (a) Rolling out voids
of the first layer of CF140 sheet once applied to the concrete block; (b) Applying
MBRACE laminate strip to prepared surface of concrete block; (e) Applying, rolling out
and removing voids from between the laminate strip and second layer of CF140. ........ 84
Figure 5.7 - Anchorage type 4 specimen geometry and material properties (WG12) .... 85
Figure 5.8 - Construction process of Type 4 Anchorage Specimen; (a) Profiling and
placement of laminate and adhesive (as per manufacturers specification) to the surface
of the concrete block; (b) Placing and rolling out voids of TYFO BCC ±45° sheet,
ensuring the direction of fibers is correct. ....................................................................... 85
Figure 5.9 - Anchorage type 5 specimen geometry and material properties (WG10 &
WG11) ............................................................................................................................. 86
Figure 5.10 - Construction process of Type 5 Anchorage Specimen; (a) Rolling out
voids of in bidirectional fabric once applied to concrete block; (b) Applied laminate
adhesive (as per manufacturers’ specification) to the surface of the bidirectional fabric
and concrete block; (c) Laminate strip ready for application of top bidirectional fabric
layer; (d) Completed anchorage specimen with two layers of TYFO BCC ±45°
bidirectional fabric sheet, positioned ±45° to the direction of the laminate strip with
laminate strip sandwiched in between. ........................................................................... 86
Figure 5.11 - Application of anchorage type 6 to proposed box girder section .............. 87
Figure 5.12 - Anchorage type 6 specimen geometry and material properties (WG8) .... 88
Figure 5.13 - Construction process of Type 6 Anchorage Specimen; The construction
sequence used for the Type 2 specimen used the following additional steps (a) Sand
back surface of cured CF140 sheet (top sheet). (b) Placing and rolling out voids of
TYFO BCC ±45° sheet, ensuring the direction of fibers is correct. ............................... 88
Figure 5.14 - Testing of type 2 (WGB3); (a) specimen ready for testing; (b) concrete
rupture at adhesive - concrete interface; (c) shear rupture of CF140 fabric at point of
wrap around; ................................................................................................................... 90
Figure 5.15 - Testing of anchor type 3 (WG6) (a) specimen ready for testing; (b)
Laminate bond failure at 1st and 2nd fabric layer interfaces; (c) 2nd layer of fabric
List of Figures
xv
rupture at base of laminate strip; (d) side view of debonded laminate strip from concrete
block. ............................................................................................................................... 91
Figure 5.16 - Testing of anchor type 4 (WG12) (a) specimen ready for testing; (b)
partial concrete-adhesive separation failure and fabric rupture (c) fabric rupture along
the ±45° fiber direction. .................................................................................................. 92
Figure 5.17 - Testing of anchor type 5 (WG10); (a) specimen ready for testing; (b) and
(c) delamination of sandwiched laminate at adhesive-concrete interface. ...................... 93
Figure 5.18 - Testing of anchor type 6 (WG8) (a) specimen ready for testing; (b) and (c)
ruptured laminate (parallel to fiber direction); (c) close up of laminate failure over
specimen free length; ...................................................................................................... 93
Figure 5.19 - Strain vs distance along Laminate; (a) Type 0 (Control) ; (b) Anchorage
Type 2 (WG3); (c) Anchorage Type 2 (WG4); (d) Anchorage Type 3 (WG5); (e)
Anchorage Type 3 (WG6); (f) Anchorage Type 3 (WG7); (g) Anchorage Type 4
(WG12); (h) Anchorage Type 5 (WG10); (i) Anchorage Type 5 (WG11); (j) Anchorage
Type 6 (WG8); ................................................................................................................ 96
Figure 5.20 - Load vs strain distribution; (a) Anchorage Type 2 (WG3); (b) Anchorage
Type 2 (WG4); .............................................................................................................. 101
Figure 5.21 - Load vs strain distribution; (a) Anchorage Type 3 (WG5); (b) Anchorage
Type 3 (WG6); Type 3 (WG7); .................................................................................... 102
Figure 5.22 - Load vs strain distribution, Anchorage Type 4 (WG12); ........................ 102
Figure 5.23 - Load vs strain distribution; (a) Anchorage Type 5 (WG10); (b) Anchorage
Type 5 (WG11); ............................................................................................................ 103
Figure 5.24 - Load vs strain distribution; (a) Anchorage Type 6 (WG8); .................... 103
Figure 5.25 - Bond-slip curves fitted with Popovics equation at bond critical regions- (a)
Type 0 (Control) ; (b) Anchorage Type 2 (WG4) ......................................................... 105
Figure 5.26 – Apparent Bond-slip curves fitted with Popovics equation at bond critical
regions (measured 125mm away from Concrete free Edge) - (a) Anchorage Type 3
(WG6); (b) Anchorage Type 4 (WG12); (c) Anchorage Type 5 (WG10); (d) Anchorage
Type 6 (WG8); .............................................................................................................. 105
Figure 5.27 – Strain of 45º Bidirectional FRP either side of laminate; Anchorage Type 4
(WG12) ......................................................................................................................... 108
Figure 5.28 – Strain of 45º Bidirectional FRP either side of laminate; Anchorage Type 5
(WG10) ......................................................................................................................... 108
List of Figures
xvi
Figure 5.29 – Strain of 45º Bidirectional FRP either side of laminate; Anchorage Type 6
(WG8) ........................................................................................................................... 109
Figure 6.1 – Stage 2, specimen summary ..................................................................... 114
Figure 6.2 – Slotted movement joints component summary ........................................ 115
Figure 6.3 – Summary of major stages of construction for stage 3 specimens; (a)
Application of first layer of bi-direction fabric; (b) Installation of FRP laminate and
creation of adhesive tapers; (c) application of final layer of bidirectional fabric and
sanding prior to application of strain gauges. ............................................................... 116
Figure 6.4 – Specimen testing rig details (a) configuration of test rig (front view); (b)
configuration of test rig (side view); (c) Photo of specimen inside testing rig ............. 117
Figure 6.5 – Speckle pattern summary; (a) speckle pattern used in stage 1; (b) improved
speckle pattern used in stage 2 ...................................................................................... 120
Figure 6.6 – Photogrammetry test set-up summary; (a) speckle patter prior to testing; (b)
CCD cameras mounted; (c) CCD cameras positioned approximately 3m away from test;
(d) typical strain data contour over entire specimen area. ............................................ 121
Figure 6.7 – Summary of pull-off testing in progress and upon completion; (a)
aluminium dolly applied prior to testing; (b) pull-off test depicting failure within
concrete; (c) pull-off test in progress; (d) pull-off test completed. ............................... 124
Figure 6.8 – Control Specimen failure summary; (a) Concrete-adhesive separation
failure (left view); (b) Back of laminate showing a combination of advesive concrete
separation failure and concrete wedge failure; (c) Concrete-adhesive separation failure
(right view) .................................................................................................................... 126
Figure 6.9 – Patch Anchor debond (Mode I); (a) front view; (b) patch anchor pull-off
failure depicting failure between saturant and the concrete .......................................... 127
Figure 6.10 – Patch Anchor debond (Mode II); (a) laminate slippage; (b) close up view
....................................................................................................................................... 127
Figure 6.11 - Load vs strain distribution; (a) Spec 0.1; (b) Spec 0.2; (c) Spec 0.3....... 131
Figure 6.12 - Load vs strain distribution; (a) Spec 1.1; (b) Spec 1.2 ............................ 132
Figure 6.13 - Load vs strain distribution; (a) Spec 2.1; (b) Spec 2.2 ............................ 132
Figure 6.14 - Load vs strain distribution; (a) Spec 3.1; (b) Spec 3.2; (c) Spec 3.3; (d)
Spec 3.4 ......................................................................................................................... 133
Figure 6.15 - Load vs strain distribution; (a) Spec 4.1, (b) Spec 4.2, (a) Spec 4.3, (b)
Spec 4.4 ......................................................................................................................... 134
List of Figures
xvii
Figure 6.16 - Strain vs distance along Laminate; (a) Spec 0.1; (b) Spec 0.2; (c) Spec 0.3
....................................................................................................................................... 135
Figure 6.17 - Strain vs distance along Laminate; (a) Spec 1.1; (b) Spec 1.2; ............... 136
Figure 6.18 - Strain vs distance along Laminate; (a) Spec 2.1; (b) Spec 2.2; ............... 136
Figure 6.19 - Strain vs distance along Laminate; (a) Spec 3.1; (b) Spec 3.2; (c) Spec 3.3;
(d) Spec 3.4; .................................................................................................................. 137
Figure 6.20 - Strain vs distance along Laminate; (a) Spec 4.1; (b) Spec 4.2; (c) Spec
4.3; (d) Spec 4.4. .......................................................................................................... 137
Figure 6.21 – Strain of 45º Bidirectional FRP either side of laminate; (a) Spec 1.1, (b)
Spec 1.2 ......................................................................................................................... 139
Figure 6.22 – Strain of 45º Bidirectional FRP either side of laminate; (a) Spec 2.1, (b)
Spec 2.2 ......................................................................................................................... 140
Figure 6.23 – Strain of 45º Bidirectional FRP either side of laminate; (a) Spec 3.1, (b)
Spec 3.2, (c) Spec 3.3, (d) Spec 3.4 ............................................................................ 141
Figure 6.24 – Strain of 45º Bidirectional FRP either side of laminate; (a) Spec 4.1, (b)
Spec 4.2, (c) Spec 4.3, (d) Spec 4.4 ............................................................................ 143
Figure 6.25 – Bond-slip curve for interface derived from experimental data, (a) Spec
0.1; (a) Spec 0.2; (a) Spec 0.3; ...................................................................................... 144
Figure 7.1 – Summary of FE model built in ATENA 3D ............................................. 148
Figure 7.2 – Calibrated shear strength interface material model and cohesion softening
law; (a) Numerical definition (Cervenka 2007); (b) shear-slip curve for interface derived
from experimental data. Where: = interfacial shear stress, = normal stress, =
friction angel, Ktt = tangential stiffness, GFI= mode 1 fracture energy, c = cohesion ... 154
Figure 7.3 – Geometry of CCIso Brick element ........................................................... 156
Figure 7.4 – FE mesh summary: (a) Type 0 (Control); (b) Type 2 (Unidirectional patch
anchor); (c) Type 5 (Bidirectional patch anchor) .......................................................... 156
Figure 7.5 – FE model boundary conditions summary. ................................................ 157
Figure 7.6 – Failure model of Control Specimen (FEM Model) depicting exaggerated
deformations .................................................................................................................. 159
Figure 7.7 – Load vs strain distribution: (a) Strain vs distance along laminate; (b) Gauge
G3; (c) Gauge G4; (d) Gauge G5; (e) Gauge G6 .......................................................... 160
Figure 7.8 – Failure model of Anchor Type 5 (FEM Model) depicting exaggerated
deformations .................................................................................................................. 162
List of Figures
xviii
Figure 7.9 - Load vs strain distribution – Type 5 (Bidirectional fabric); (a) Strain vs
Distance along FRP laminate; (b) Gauge G3; (c) Gauge G4; (d) Gauge G5; (e) Gauge
G6. ................................................................................................................................. 163
Figure 7.10 - Type 5 (45º Bidirectional FRP) – Strain parallel to the fibers of the patch
anchor (±45º) vs distance away from centre of laminate (mm) ................................... 165
Figure 7.11 – Failure model of Anchorage Type 3 (FEM Model) depicting exaggerated
deformations.................................................................................................................. 166
Figure 7.12 - Load vs strain distribution, Type 2 (Unidirectional fabric); (a) Gauge G3;
(b) Gauge G4; (c) Gauge G5; (d) Gauge G6 ................................................................. 167
Figure 7.13 - Comparison of load-strain curves predicted by the models with different
mesh sizes ..................................................................................................................... 170
Figure 7.14 - Comparison of bond-slip curves predicted by the models with different
mesh sizes ..................................................................................................................... 171
Figure 7.15 - Comparison of load-strain curves predicted by the models with different
fracture energy .............................................................................................................. 173
Figure 7.16 - Comparison of load-slip curves predicted by the models with different
fracture energy .............................................................................................................. 174
Figure 7.17 - Comparison of load-strain curves predicted by the models with different
adhesive stiffness .......................................................................................................... 176
Figure 7.18 - Comparison of load-strain curves predicted by the models with different
concrete strengths .......................................................................................................... 178
Figure 7.19 - Anchorage Type 5 parametric study – Concrete strength vs max laminate
strain prior to de-bond. .................................................................................................. 179
Figure 8.1 – Summary of finite element model components. ....................................... 181
Figure 8.2 - Typical interface model behaviour in shear with cohesion softening law; (a)
Numerical definition (Cervenka 2007); (b) shear-slip curve for interface derived from
experimental data. Where: = interfacial shear stress, = normal stress, = friction
angel, Ktt = tangential stiffness, GFI= mode 1 fracture energy...................................... 183
Figure 8.3 – Failure model of Control Specimen (FEM Model) depicting exaggerated
deformations.................................................................................................................. 185
Figure 8.4 – Failure model of Anchor Type 1 (FEM Model) depicting exaggerated
deformations.................................................................................................................. 187
List of Figures
xix
Figure 8.5 – Failure model of Anchor Type 2 (FEM Model) depicting exaggerated
deformations .................................................................................................................. 188
Figure 8.6 – Failure model of Anchor Type 3 (FEM Model) depicting exaggerated
deformations .................................................................................................................. 189
Figure 8.7 – Failure model of Anchor Type 4 (FEM Model) depicting exaggerated
deformations .................................................................................................................. 190
Figure 8.8 – Type 0.1 Strain distribution Summary: (a) Strain vs Distance along FRP
laminate; (b) Gauge G2; (c) Gauge G3; (d) Gauge G4; (e) Gauge G5; (f) Gauge G6 ; (g)
Gauge G7 ...................................................................................................................... 194
Figure 8.9 – Type 1.2 Strain distribution Summary: (a) Strain vs Distance along FRP
laminate; (b) Gauge G2; (c) Gauge G3; (d) Gauge G4; (e) Gauge G5; (f) Gauge G6 ; (g)
Gauge G7 ...................................................................................................................... 195
Figure 8.10 – Type 2.2 Strain distribution Summary: (a) Strain vs Distance along FRP
laminate; (b) Gauge G2; (c) Gauge G3; (d) Gauge G4; (e) Gauge G5; (f) Gauge G6 ; (g)
Gauge G7. ..................................................................................................................... 196
Figure 8.11 – Type 3.4 Strain distribution Summary: (a) Strain vs Distance along FRP
laminate; (b) Gauge G2; (c) Gauge G3; (d) Gauge G4; (e) Gauge G5; (f) Gauge G6 ; (g)
Gauge G7. ..................................................................................................................... 197
Figure 8.12 – Type 4.4 Strain distribution Summary: (a) Strain vs Distance along FRP
laminate; (b) Gauge G2; (c) Gauge G3; (d) Gauge G4; (e) Gauge G5; (f) Gauge G6 ; (g)
Gauge G7. ..................................................................................................................... 198
Figure 8.13 – Strain of 45º Bidirectional FRP either side of laminate; (a) Spec 1.2, (b)
Spec 2.2, (c) Spec 3.4, (d) Spec 4.4 .......................................................................... 200
Figure 8.14 - Comparison of load-strain curves predicted by the models with different
concrete strengths .......................................................................................................... 202
Figure 8.15 - Anchorage Type 5 parametric study – Concrete strength vs max laminate
strain prior to de-bond. .................................................................................................. 202
Figure 9.1 – Summary of parametric study results conducted on concrete strength and
the maximum FRP strain reached prior to debond........................................................ 206
Figure 9.2 – Typical strain overlay in bidirectional fibers resulting from superposition of
strain between two adjacent laminates. ......................................................................... 209
List of Tables
xx
LIST OF TABLES
Table 2.1 - FRP anchorage summary for flexurally strengthened members .................. 25
Table 2.2 - CFRP Shear Anchorage devices summary ................................................... 37
Table 3.1 – Summary of proposed bond slip models for FRP-to-concrete joints ........... 49
Table 3.2 – Summary of proposed bond strength models for FRP-to-concrete joints .... 54
Table 4.1 - Summary of test specimens constructed in experimental program .............. 57
Table 4.2 - FRP Properties data ...................................................................................... 59
Table 4.3 - Adhesives and Saturant Properties data ........................................................ 60
Table 4.4 - Adhesion test results on TYFO BCC bidirectional fabric and MBRACE
laminate strip. .................................................................................................................. 66
Table 4.5 - Load/Elongation results summary (WG1, WG2 & WG9) ........................... 72
Table 4.6 - Max Bond stress and corresponding slip results summary (WG1, WG2 &
WG9) at location 125mm away from concrete free edge. .............................................. 74
Table 4.7 - Maximum FRP elongations and corresponding effective FRP strains and
utilisation percentiles ...................................................................................................... 76
Table 5.1 – Summary of test specimens constructed in experimental program ............. 80
Table 5.2 – Maximum FRP elongations and corresponding effective FRP strains and
utilisation percentiles (types 0, 2-6) ................................................................................ 94
Table 5.3 – Bond stress and corresponding slip results summary (type 0, 2-6) ........... 104
Table 6.1 – Summary of test specimens constructed in experimental program ........... 113
Table 6.2 – Adhesives, Saturant and Primer data ......................................................... 118
Table 6.3 –FRP Properties data..................................................................................... 118
Table 6.4 – Concrete Mechanical Properties - Compressive cylinder results summary;
....................................................................................................................................... 122
Table 6.5 – Concrete Mechanical Properties - Pull-off test results summary ............... 124
Table 6.6 – Adhesive Mechanical Properties - Tensile dumbbell results ..................... 125
Table 6.7 – Results summary ........................................................................................ 128
Table 6.8 – Bond stress and corresponding slip results summary (type 1-4) ............... 143
Table 6.9 – Bond stress and corresponding slip results summary (type 1-4) ............... 145
Table 7.1 – Summary of input parameters used in non-linear concrete model ............ 151
Table 7.2 – Parameters used to define interface material model .................................. 154
List of Tables
xxi
Table 7.3 – Summary of maximum loads and FRP strains reached prior to debonding
derived from experimental data (types 0, 2 & 5) .......................................................... 158
Table 7.4 - Mesh size variations ................................................................................... 168
Table 7.5 - Concrete fracture energy variations ............................................................ 172
Table 7.6 - Adhesive modulus variations ...................................................................... 175
Table 7.7 – Summary of material properties used to evaluate sensitivity to concrete
strength .......................................................................................................................... 177
Table 8.1 - Concrete material model parameters used in numerical model .................. 181
Table 8.2 - Interface material model parameters used in numerical model .................. 184
Table 8.3 – Results summary ........................................................................................ 191
Table 8.4 – Summary of material properties used to evaluate sensitivity to concrete
strength .......................................................................................................................... 201
Table 9.1 - Summary of strength prediction models compared with FRP-to-Concrete
joints .............................................................................................................................. 205
Table 9.2– Summary of experimental and numerical predictions, verified with the
proposed anchorage strength mode ............................................................................... 213
Notation and Abbreviation
xxii
NOTATION AND ABBREVIATION
Constant in Popovics’ equation
AR ARAMIS image correlation photogrammetry
Cross-sectional area of a FRP composite
ASF Adhesive separation failure
Factor to account for the width ratio between the FRP and
concrete in bond slip models
Coefficient in bond strength model proposed by Chen and
Teng 2001 to account for s reduced FRP bond length
Coefficient in bond strength model proposed by Chen and
Teng 2001 to account for the width ratio between FRP
and concrete
Modification factor to account for adhesive stiffness, fiber
stiffness and fiber thickness
Width of a concrete member
Width of an adhesive layer
Width of the bonded FRP plate
Width of a FRP composite
Width of a web
c Cohesion coefficient and cement content in fracture
energy formula
Constant determined from a regression analysis of FRP
pull test
CFS Cover separation failure
CDC Critical diagonal crack debonding
Maximum aggregate size
Effective depth of FRP shear reinforcement
Modulus of elasticity of a FRP composite
Modulus of elasticity of adhesive
Modulus of elasticity of concrete
Notation and Abbreviation
xxiii
Modulus of elasticity of a FRP composite
Modulus of elasticity of tension steel reinforcement
The FRP maximum effective strain is used to determine
the anchorage effectiveness factor
Design rupture strain of FRP reinforcement
FRP strain value at ith +1 load increment used to calculate
bond-slip data
FRP strain value at ith load increment used to calculate
bond-slip data
Effective strain level in FRP reinforcement attained at failure,
Concrete surface tensile strength determined in a pull-off
test
f’cm Mean tensile strength of concrete
fct Characteristic principle tensile strength of concrete
FR Fabric Rupture
Characteristic compressive cylinder strength of concrete
GFI Specific mode-I fracture energy of concrete
GFII Specific mode-II fracture energy of concrete
Shear stiffness of FRP reinforcement;
Shear modulus of an adhesive
IC Intermediate crack induced debonding
Second moment of area of the FRP plate
K Ratio between adhesive shear modulus and thickness used
to determine mode I fracture energy
Factor used to determine effective FRP strain for shear
strengthening
Anchorage effectiveness factor for shear strengthened
members
Modification factor applied to v to account for concrete
strength
Modification factor applied to v to account for wrapping
scheme
Knn normal stiffness of interface in finite element model
Notation and Abbreviation
xxiv
Ktt tangential stiffness of interface in finite element model
ratio between shear modulus of adhesive and adhesive
thickness
kp Coefficient in bond strength model proposed by Neubauer
1997 to account for the width ratio between FRP and
concrete
Coefficient to account for the width ratio between FRP
and concrete used to determine model 1 fracture energy of
concrete
Anchorage effectiveness factor for flexurally strengthened
members
L Bonded length
Effective bond length
Length of FRP Patch anchor
Distance between two monitoring points used to calculate
bond slip data
LR Laminate Rupture
LS Fabric right ride of laminate
n Number of plies of FRP composite
Number of plies of FRP composite
PASF Partial Adhesive Separation Failure
Maximum anchorage strength of Patch Anchored joint
Bond strength of a joint
r1 Modification factor for concrete compressive strength
used to calculate the strength of patch anchored joints.
r2 Modification factor for FRP laminate width used to
calculate the strength of patch anchored joints. r3 Modification factor for FRP spacing used to calculate the
strength of patch anchored joints. r4 Modification factor for patch anchor length used to
calculate the strength of patch anchored joints. RS Fabric right side of laminate
SG Strain gauge
Notation and Abbreviation
xxv
FRP slip value at ith load increment used to calculate bond-
slip data
FRP slip value at ith - 1 load increment used to calculate
bond-slip data
s FRP Slip
Slip corresponding to peak bond stress along the bond-
slip curve
Slip at which the bond stress reaches its maximum value
(bond-slip curve context)
Maximum shear strength of an interface
Thickness of an adhesive layer
Shear stress
Thickness of a FRP composite
Thickness of a FRP composite
Maximum shear stress prior to debonding
V3D Vic3D Image correlation photogrammetry
V Displacement
Stress normal to interface in Interface element definition
Coefficient in bond strength model proposed by Neubauer
1997 to account for a reduced FRP bond length
Factor relating to aggregate type used to calculate model I
fracture energy
Chapter 1 - Introduction
1
CHAPTER 1 – INTRODUCTION
1.1 General background
The present expansion of civil infrastructure to meet the demands of escalating
population growth have resulted in introduction of higher bridge loading specifications
implemented in many countries. Environmental degradation and stricter bridge design
and evaluation specifications (CAN/CSA-S6 1988; AASHTO 1994; ACI 440.2R-08
2008) have contributed to the increasing number of bridge retrofit projects currently
underway. Due to the need of structural rehabilitation of buildings, bridges as well as
other infrastructure, any new rehabilitation technologies have been introduced in recent
years. Using carbon fiber reinforced polymers (FRP’s) has become a popular method to
strengthen existing buildings and bridge structures which are being subjected to higher
bending, shear and torsional forces. A good example of this is the West Gate Bridge in
Melbourne, Australia, which is currently the world’s largest application of FRP
composites used to strengthen a structure (Hii and Al-Mahaidi 2006).
Strengthening using FRP’s has been proven to provide a suitable and cost effective
solution to strengthening of existing concrete structures due to: deterioration, increased
loads and changes in usage. The material has several well documented advantages over
traditional strengthening materials such as: light weight, high tensile strength,
durability, ease of installation and unobtrusiveness (Khalifa, Belarbi et al. 2000).
Application of FRP has been found to be more cost effective than traditional
strengthening methods and can negate the need to replace the original structure. The
cost effectiveness of application of FRP composites is further increased by the ongoing
savings associated with lack of maintenance. As a result, FRP strengthening is enjoying
a great deal of popularity in the construction industry (Rizkalla and Nanni 2003).
FRP laminates or sheets are typically applied to structural members as externally
bonded reinforcement using high strength adhesives or as near surface mounted
reinforcement. Fibers are bonded to the concrete only after sufficient surface
preparation consisting of: grit blasting, water jetting (to expose aggregate) and
application of a suitable primer. Previous experimental studies have shown that the
effectiveness of the strengthening technique is largely governed by the bond behaviour
Chapter 1 - Introduction
2
between the FRP and concrete. As a result, failure of FRP-to-concrete joints typically
occurs by debonding of the FRP from the concrete substrate.
1.2 Research problem and aims
Research has demonstrated that the effectiveness of the FRP when applied to concrete
members is largely governed by the strength of the bond between the FRP and the
concrete. As a result, failure of strengthened members is usually a result of FRP
debonding from the concrete substrate. FRP design guidelines such as: (Fib: Task
Group 9.3 2001; Concrete Society Technical Report No. 55 2004; ACI 440.2R-08 2008)
acknowledge many modes of deboning failure for flexural and shear strengthened
members such as: The modes are summarized as concrete crushing, FRP rupture, shear
failure, concrete cover separation failure, plate end interfacial debonding, intermediate
flexural or flexural-shear crack-induced interfacial debonding (otherwise known as
IC debonding) and shear-induced debonding [also referred to as critical diagonal crack
(CDC) debonding]. Such premature failure modes are prevented at design stage by
limiting the stresses within the FRP-to-concrete interface. Typically, less than half of
the fiber rupture strain is utilized in design to prevent debond and in cases where a high
degree of strengthening is required (typical of many bridge applications where thicker
fibers are required).
Poor fiber utilization levels are particularly observed in shear strengthening
applications, where effective FRP design strains are typically only 10-25% of the fiber
rupture strain. One of the primary methods to improve the bond performance of the FRP
material to the concrete surface is by the provision of anchorage which could include:
FRP U-jackets, FRP anchors (also known as spike anchors amongst other names), patch
anchors (utilising unidirectional and bidirectional fabrics), nailed metal plates (also
known as hybrid bonding), near-surface mounted rods, mechanical fastening, concrete
embedment and mechanical substrate strengthening. Despite the wide range of anchors
investigated in the literature, research in the field remains in its infancy due to a lack of
experimental data and an absence of design guidelines. Furthermore, many of the
proposed anchorage systems have been unappealing due to their high installation costs
Chapter 1 - Introduction
3
and need for mechanical fasteners which have the potential to damage the existing
structure.
The general aim of this dissertation is, to develop a new type of anchorage for EB FRP
which satisfies the criteria of: prevention of premature FRP debonding, non-
destructiveness (absence of mechanical fasteners), ease of installation, high durability,
low maintenance and cost efficiency. The specific aims are:
Summarize and assess the efficiency of existing anchorage systems, their
advantages, disadvantages and scope for application.
To summarise the various theoretical models for predicting the anchorage
strength and bond behaviour of FRP-to-concrete joints.
Conduct an experimental investigation into the use of mechanical substrate
strengthening to increase the bond strength between the FRP and concrete
medium.
Study the effect of unidirectional fibers, orientated parallel and perpendicular to
the direction of loading and their effect on anchorage strength via an
experimental program.
Conduct an experimental investigation to a novel anchorage system, herein
named ‘patch anchor’ consisting of (±45º) bidirectional fibers applied to the
ends of FRP laminates, via a series of full scale FRP-to-concrete shear tests.
Perform further experimental studies into the size effect of (±45º) bidirectional
fiber patch anchors and investigate alternative laminate properties on anchorage
performance.
Develop 3D, non-linear, calibrated, finite element models for patch anchor
specimens and conduct parametric studies once verified with the experimental
data.
Develop theoretical design formulations for (±45º) bidirectional fiber patch
anchors, taking into account parameters such as: patch anchor size, FRP grade,
thickness, width and concrete compressive strength.
Chapter 1 - Introduction
4
1.3 Thesis outline
This thesis is divided into three main parts. Part I presents an overview of the previous
research relevant to the study. Chapter 2 presents a state-of-the-art literature review of
previous experimental investigations conducted in the field of FRP anchorage systems.
A database consisting of collected experimental results on the various anchorage types
is constructed resulting in an anchorage efficiency factor assigned to each anchor for
detailed evaluation. Chapter 3 summarises the important information associated with
FRP-to-concrete bond behaviour, anchorage strength prediction, bond slip models and
important parameters, where an understanding is needed, in order to proceed with the
experimental and finite element works to follow.
Part II presents the main body of experimental work undertaken in the study. Chapter 4
reports an experimental investigation to the use of a mechanically strengthened
substrate to enhance the bond performance of FRP to concrete. Chapter 5 commences
the first stage of the experimental investigation into the use of unidirectional and (±45º)
bidirectional fiber patch anchors applied to the end of FRP laminates to enhance the
strength of the joints. A total of 10 full scale specimens were constructed and the
various anchorage joint configurations were tested in direct shear. Chapter 6 presents
stage 2 of the experimental study, the focus of which is the investigation into the size
effect of (±45º) bidirectional fiber patch anchors and results in the testing of a further 16
large scale specimens.
The development of the finite element and theoretical design models to describe the
behaviour of the patch anchored specimens under direct shear is presented in Part III,
which consists of three chapters. Chapters 7 and 8 present the finite element simulations
and parametric studies conducted on the specimens tested in stages 1 and 2 of the
experimental programs. The development and verification of a new theoretical model
for predicting the capacity of FRP-to-concrete patch anchored joints is presented in
chapter 9. Finally, in chapter 10, major conclusions from this research are presented
with recommendations for future studies.
Chapter 2 – Literature Review of FRP Anchorage Systems in concrete Infrastructure
5
2 CHAPTER 2 - LITERATURE REVIEW OF FRP ANCHORAGE SYSTEMS IN CONCRETE INFRASTRUCTURE
2.1 Introduction
The retrofitting of existing reinforced concrete (RC) structures has become necessary
due to environmental degradation, changes in usage, and heavier loading conditions. In
the forefront of retrofitting technology is the use of advanced fiber-reinforced polymer
(FRP) composites applied to structural members as externally bonded reinforcement
(Bank 2006; Hollaway and Teng 2008; Karbhari and Abanilla 2007). The suitability of
this material when compared, for example, to structural steel is largely due to its light
weight, superior tensile strength, and its resistance to corrosion. These FRP materials
are typically applied to the concrete surface using epoxy resin after adequate surface
preparation of the concrete, typically involving sandblasting, water jetting, and the
application of a suitable primer. Once applied, up to seven days of curing is typically
required to achieve the full bond strength of the system (Hag-Elsafi et al. 2001).
However, FRP solutions are not without their inherent shortcomings. For instance, it is
widely recognized that failure of RC structures retrofitted with FRP almost always
occurs by debonding of the FRP from the concrete substrate. To prevent this type of
failure, national standards and design guidelines impose strict limitations on the
allowable strain level in the FRP which may be safely utilized in design. To achieve
acceptable levels of FRP-to-concrete contact bond stress, allowable strains are lower in
cases where a higher degree of strengthening is required and can be as low as 10–25%
of the material rupture strain. Low levels of efficiency are often the result of using
higher modulus fibers and multiple layers of FRP. In practice these limitations result in
severe underutilization of the FRP material properties. Anchorage of the FRP is one
means to significantly improve the efficiency of FRP systems and hence provide a
solution to these shortcomings. Extensive research has been undertaken to understand
the mechanisms of FRP application and failure and has resulted in design guidelines
being published all around the world within the last decade [e.g., International
Federation for Structural Concrete (fib) 2001; Japan Society of Civil Engineers (JSCE)
2001; Concrete Society 2004; American Concrete Institute (ACI) 2008; Oehlers et al.
Chapter 2 – Literature Review of FRP Anchorage Systems in concrete Infrastructure
6
2008]. It is understood that the bond strength of FRP materials can be improved when
sufficient anchorage is provided and such provisions have been acknowledged to delay
or prevent the critical mode of FRP debonding failure (Galal and Mofidi 2010). In
addition, anchorage devices can be essential to transfer the stress from one structural
component to another where application is limited by the geometrical configuration. A
popular example is the shear strengthening of T-shaped sections (Ceroni et al. 2008).
The primary obstacle presently preventing the widespread use of FRP anchorage
measures is that no rational and reliable design rules currently exist. As a result, FRP
design guidelines stipulate that the practical implementation of anchorage devices
should be substantiated by representative experimental testing (ACI 2008). However,
the guidelines do not specify the types of testing procedures that are considered
adequate (Grelle and Sneed 2011). The repercussions of time and budget constraints on
small and large scale industrial projects mean that such testing is rarely carried out in
practice. As a result, the potential benefits of FRP anchorages have typically been
superseded by more conservative strengthening approaches such as section
enlargement or column insertion.
Although anchorage devices applied to the ends of FRP reinforcements have been
tested by many researchers, the results have been limited by case dependency with
relatively small sample sizes being employed for each study. This chapter provides a
review of representative experimental studies conducted on the major anchorage
concepts by drawing upon a wide selection of publications. The chapter assumes a
largely qualitative style by physically explaining each anchor concept with the aid of
appropriate diagrams. Information about typical experimental investigations
undertaken on each anchor type and descriptions of behaviour and failure are given.
Databases are also assembled from available test results and efficiency factors are
calculated for each anchor concept. Such calculations represent the quantitative aspect
of the paper. While it is recognized that anchorages can be of benefit to a variety of
FRP-strengthened elements such as connections, wall, and beams members, emphasis
has been given in this paper to flexural members strengthened in flexure and shear
because these constitute the most common strengthening situations. Finally, the terms
retrofitting and strengthening are used interchangeably throughout the document.
Chapter 2 – Literature Review of FRP Anchorage Systems in concrete Infrastructure
7
2.2 Mechanisms of FRP failure and debonding for flexurally strengthened members
To date, several failure modes for RC beams strengthened in flexure with FRP plates
have been identified from experimental investigations and these are shown in Figure
2.1. The modes are summarized as (1) concrete crushing, (2) FRP rupture, (3) shear
failure, (4) concrete cover separation failure (Yao and Teng 2007), (5) plate end
interfacial debonding (Leung and Yang 2006), (6) intermediate flexural or flexural-
shear crack-induced interfacial debonding (otherwise known as IC debonding) Teng
et al. 2003; Ombres 2010), and (7) shear-induced debonding [also referred to as critical
diagonal crack (CDC) debonding] (Oehlers and Seracino 2004; Wang and Zhang
2008). Modes 4 to 7 are all premature debonding failures. Of these, modes 4 and 5
initiate at or near the plate end, while modes 6 and 7 initiate away from the plate end.
In addition, modes 5 and 6, and sometimes mode 7, occur at the FRP-to-concrete
interface (in the concrete), while modes 4 and 7 can occur predominantly at the internal
steel reinforcement level. Detailed accounts of all failure modes are provided elsewhere
(Hollaway and Teng 2008).
Figure 2.1 - Types of CFRP debonding (adapted from Pham and Al-Mahaidi 2004)
Many factors control the likelihood of a particular debonding failure mode, including
(1) the level of internal steel reinforcement, (2) the distance between a plate end and the
Chapter 2 – Literature Review of FRP Anchorage Systems in concrete Infrastructure
8
adjacent beam support (plate end distance), (3) FRP plate length, width, thickness, and
modulus of elasticity, (4) shear-to-moment interaction, (5) concrete strength, and (6)
section geometry (Teng and Yao 2007). Observations suggest that as the plate end
moves further away from the support, cover separation failure becomes the controlling
mode, whereas IC debonding governs when the distance between the plate end and
support is relatively small (Yao and Teng 2007). In addition, the probability of
debonding initiating near the plate end has been found to be the highest when the ratio
of maximum shear force to bending moment is high, such as the higher peeling stresses
generated at the ends of the external plate. Therefore, slender beams with high shear
span/depth ratios do not present a need for plate end anchorage because failures are
initiated in regions of high bending moment well away from the plate ends (e.g.,
Garden and Hollaway 1998). These are just some of many qualitative observations to
be found in the published literature.
2.3 Anchorage devices for FRP reinforcement used to strengthen members in flexure
Three general categories of anchorage type have been investigated to date to prevent
debonding in RC members strengthened in flexure with FRP, namely:
(a) U-jacket anchors (Smith and Teng 2003; Al-Amery and Al-Mahaidi 2006; Pham
and Al-Mahaidi 2006; Yalim, Kalayci et al. 2008),
(b) Mechanically fastened metallic anchors (Garden and Hollaway 1998; Spadea et
al. 1998; Jensen et al. 1999; Duthinh and Starnes 2001; Wu and Huang 2008), and
(c) FRP anchors (Lam and Teng 2001; Eshwar et al. 2005; Micelli et al. 2010; Smith
2010; Zhang and Smith 2012a, b; Zhang et al. 2012).
2.3.1 FRP U-jacket anchors
FRP U-jacket anchors involve the application of unidirectional or bidirectional fiber
to the ends of flexural FRP reinforcement (Figure. 2) to prevent or delay debonding
initiating from the plate end. U-jackets can also be placed along the length of the
member to prevent or delay debonding initiating away from the plate end. The ultimate
Chapter 2 – Literature Review of FRP Anchorage Systems in concrete Infrastructure
9
function of a U-jacket is to provide the confinement necessary to resist the tensile
peeling stresses and longitudinal crack propagation at fiber termination points or
intermediate cracks. Khan and Ayub (2010) investigated anchorage heights ranging
from 100–200 mm and suggested that U-shaped anchorages were effective irrespective
of their height. The study deter- mined that 100 mm partial-height U-wraps delivered
the same effectiveness as full-height U-wraps because in both cases failure was by
concrete crushing. Because concrete crushing was observed for the shorter length
jackets, the true potential of full-height jackets could not be utilized.
Debonding failure modes can change due to the addition of FRP U-jackets. For
example, Smith and Teng (2003) showed that with the addition of plate-end U-jackets,
the critical debonding failure mode could be shifted from concrete cover separation to
IC de- bonding. Therefore, in an effort to prevent failure by IC debonding, the
placement of U-jackets throughout the span or in the flexural- shear zones (at certain
spacing’s) has been investigated by several researchers to date (Al-Amery and Al-
Mahaidi 2006; Khan and Ayub 2010; Pham and Al-Mahaidi 2006; Yalim et al. 2008).
Although lacking in material efficiency, this method has been proven to result in FRP
rupture. Such an arrangement of U-jackets is also used for shear strengthening
applications. Selected studies are summarized in the following:
IC debonding in beams retrofitted with U-jacket anchors was re- ported by Pham and
Al-Mahaidi (2006). The experimental program comprised 260 × 140 mm RC beams
tested under three-and four- point bending. Anchorages encompassing unidirectional
fibers of 209 GPa modulus were placed at the carbon FRP (CFRP) plate ends or at a
spacing of 180 mm within the shear zone. Each jacket com- prised two plies of fabric
that was 0.175 mm thick and 50 mm wide, which was bonded to the sides and the soffit
of the concrete beam to form a U-shape. While the end U-jacket proved to be effective
in limiting both forms of end debonding, i.e., end cover separation failure and end
interfacial debonding, the critical failure mode was seen to shift to intermediate-span
debonding at a higher load, and it often occurred together with rupture of the end U-
jacket. Such behaviour was also observed in Smith and Teng’s (2003) study. The
rupture was due to a sliding action of the CFRP reinforcement underneath the U-jacket,
causing bending of the jacket legs near the soffit. The experiments also confirmed that
Chapter 2 – Literature Review of FRP Anchorage Systems in concrete Infrastructure
10
the placement of U-jackets in the shear span at certain spacing can postpone the
occurrence of IC de- bonding. The inclusion of U-jackets in the shear zone had the dual
benefits of resisting the opening of flexural-shear cracks and improving the CFRP-to-
concrete bond strength by the increased level of confinement underneath the U-jacket.
To further understand the confining action of FRP U-jacket anchors, the vertical strain
distribution within the vertical FRP legs was investigated by Sawada et al. (2003). The
strains reported reached values of 3;000 in the cover region of the concrete and at a
load level expected to produce debonding. Further load application resulted in 6;000
being recorded at the maximum loading point. This is indicative that the CFRP U-jacket
was resisting the stresses that typically result in cover separation failure.
Further research conducted by Al-Amery and Al-Mahaidi (2006) determined that the
use of the CFRP U-jackets at 200 mm spacing along the length of the beam reduced the
interfacial slip between the CFRP flexural fiber and the concrete section by up to one-
tenth. In this study, the U-jackets lead to the full utilization of the CFRP flexural tensile
capacity. The results demonstrated an increase in flexural strength of up to 95% when
using CFRP U-jackets to anchor the CFRP fiber. However, when using conventional
CFRP fibers alone, an increase of only 15% was achieved. Yalim et al. (2008) also
conducted investigations on the effects of U-jacket configurations placed throughout the
span as opposed to only the plate ends. A total of 26 beams were tested in 3-point
loading with 4, 7, 11, and continuous U-jacket arrangements. The study utilized FRP U-
jackets to anchor both FRP laminates (modulus of elasticity of 131 GPa) and FRP
sheets (modulus of elasticity of 70.6 GPa). In addition, three alternative surface profiles
were investigated: smooth, intermediate, and rough. However, each surface profile was
not appropriately defined (except by broad definition) and as a result, the categorization
is not an appropriate definition of surface roughness. The use of four U-jackets at the
FRP ends was successful in preventing the end interfacial debonding failure that was
observed in unanchored specimens, and failure was shifted to IC debonding, confirming
the findings of earlier researchers. The beams with seven jackets failed in the same way
at a higher load together with U-jacket debonding. Specimens with eleven jackets and
full continuous jackets failed by rupture of FRP. Although the strain utilization levels
and ultimate load capacity were improved with the addition of U-jackets throughout the
Chapter 2 – Literature Review of FRP Anchorage Systems in concrete Infrastructure
11
span, it was found that a higher level of anchorage improved the ductility more than it
did the strength. However, the ductility measurements were solely based on the
maximum vertical deflection for the beams prior to failure. Ductility can be defined as
the RC beam’s ability to deform under tensile stress and can be determined by
monitoring deflection, beam curvature, or strain in the tensile reinforcement.
Monitoring beam deflection may be indicative of ductile behaviour, but the method fails
to consider deformability in terms of beam curvature and cracking (measured from
tensile reinforcement strain). In addition, most FRP design guidelines check strain of
the tensile reinforcement to ensure ductility. Although the benefit of U-jacket anchors in
flexural retrofitting applications is evident, the provision of U-jackets throughout the
span to prevent the mechanisms of plate end and IC debonding may not be a materially
efficient method to improve the efficiency of FRP strengthening applications because
additional material is required to reach a given strength (Orton et al. 2008).
2.3.2 Inclined U-jacket orientations
Promising results have been achieved based on the limited research conducted on
inclined U-jackets at FRP ends only (refer figure 2.2). Published findings indicate that
in addition to preventing the two mechanisms of end span debond, inclined anchors
readily shift the critical failure mode to concrete crushing or FRP rupture (Duthinh &
Starnes 2001; Pornpongsaroj & Pimanmas 2003; Sagawa 2001).
Figure 2.2 – U-shape anchoring method 45 degree
Chapter 2 – Literature Review of FRP Anchorage Systems in concrete Infrastructure
12
The effects of alternative U-jacket orientations, including perpendicular, inclined, and
X-shaped U-jacket anchors, were investigated by Pimanmas and Pornpongsaroj (2004).
In this study, 220 mm deep and 120 mm wide RC beams were tested under four-point
bending. Beams were retrofitted with 1.2 mm thick and 100 mm wide plates for flexural
strengthening with a 150 GPa modulus of elasticity. The plates were anchored at the
plate ends with 0.11 mm thick carbon fiber sheets over a width of 300 mm. Anchorages
consisted of the application of a single ply of CFRP with 230 GPa material stiffness.
The study investigated two plate-end termination lengths: 200 mm and 420 mm away
from the supports, which failed by IC debonding and end cover separation failure,
respectively, where no anchorage was provided.
Of the numerous anchor configurations tested, it was found that U-jackets placed at the
FRP plate-end locations 200 mm from sup- ports failed by premature concrete crushing
and intermediate span debonding, while U-jackets placed 420 mm away from supports
failed by premature concrete crushing and concrete cover separation failure. The
influence of end termination distance on end de- bonding failure is consistent with
current debonding models (Smith and Teng 2002; Smith and Teng 2003). Inclined and
X-shaped anchor arrangements all failed by concrete crushing. Interestingly, the authors
point out that the CFRP plate experienced the highest confinement near the side faces of
the beam and less restraint in the central zone. This implies that U-jacket anchorages
lose effective- ness with increasing beam width. Although the authors concluded that
the inclined and X-shaped anchors successfully prevented both forms of plate end and
IC debonding, premature concrete crushing failure prevented the occurrence of FRP
rupture, masking the full potential of the anchorages from being realized.
Duthinh and Starnes (2001) also confirmed that concrete crushing was the controlling
failure mode in two out of the three specimens that they tested, and the other mode was
a combination of U-jacket rupture and intermediate flexural-shear crack debonding. The
laboratory program comprised 2–6 layers of 200 mm wide CFRP jackets placed
diagonally on each plate end. The inclined fibers effectively prevented cover separation
failure at the plate ends. It was found that two and six layers of jacket anchored the
carbon plate to strain levels of 8,260 and 11,000 , respectively, without slippage. The
above research demonstrates the clear advantages of using inclined U-jackets as
Chapter 2 – Literature Review of FRP Anchorage Systems in concrete Infrastructure
13
opposed to perpendicular orientations at the CFRP plate ends. In addition to the jackets
providing confinement, an improvement of bonding and resistance to the opening of
longitudinal cover separation cracks, the inclined fibers were seen to delay the
occurrence of IC debonding. This may be due to a reduction of interfacial longitudinal
shear stresses in the shear-flexural zones and the resulting energy transfer to the jacket
anchors via an induced strut-and-tie action resulting from the inclined fibers. The
benefits of inclined fibers were also noted by Sagawa et al. (2001).
In addition to the prevention of debonding failure, Smith and Teng (2003) showed that
the use of U-jackets can also enhance ductility. This was confirmed by Buyle-Bodin
(2004), who investigated several FRP anchorage devices to prevent concrete cover
separation failure. The experimental program involved five beams, each 3,000 mm long
with a rectangular cross-section 150 mm wide and 300 mm deep. Both perpendicular
and laterally inclined CFRP shear jackets were used to restrain the ends of the CFRP
flexural plate at 130–200 mm spacings. Ductility was measured as either deflection
ductility or curvature ductility. Deflection ductility was defined as the ratio of ultimate
midspan deflection to yield midspan deflection, whereas curvature ductility was
considered in a similar fashion but utilized the midspan curvature values. Although all
specimens strengthened with both perpendicular and inclined shear jackets exhibited
greater load-carrying capacity, deflections, and ductility, it was found that perpendicular
orientations of U-jacket anchors provided the most noticeable improvement, with
increases in curvature ductility of 45% and 24% for deflection ductility. The
improvements were less obvious in the inclined U-jacket anchors. This may be due to
the higher post cracking stiffness exhibited due to the inclined U-jacket anchors. Strain
in the tensile reinforcement is usually the most common measure of ductility utilized by
FRP design guidelines such as ACI 440.2R-08 (2008). It may be more beneficial for
future researchers to measure the tensile reinforcement strain to quantify ductility
performance.
2.3.3 Prestressed U-jackets
Prestressed U-jackets are a method of anchorage on which little research has been
conducted. The advantages of prestressing stem from the increased level of confinement
Chapter 2 – Literature Review of FRP Anchorage Systems in concrete Infrastructure
14
and the higher shear resistance provided by the prestressed U-jackets. In practical
applications, prestressing was introduced onto the sides of the CFRP U-jackets by Pham
and Al-Mahaidi (2006) by introducing a gap between the jacket and the concrete soffit,
as presented in figure 2.3.
Figure 2.3 – Prestressing system for FRP ligatures (modified from Pham and Al-Mahaidi 2006)
A prestressing strain of 500 was introduced into the jacket sides by inserting wedges
into a preformed gap. Beams with pre- stressed jackets showed no evidence of slippage
in the anchorage zone at failure. This was attributed to an increase in concrete shear
capacity in the anchorage zone as a result of the compressive stress induced by the U-
jackets. The legs of the prestressed U-jackets did not rupture, but failed through a
combination of IC debonding and debonding of the end jacket. Only a slight
improvement of approximately 5% in the ultimate capacity was recorded due to
prestressing. Debonding of the U-jackets suggests that a more robust form of anchorage
is required to anchor the ends of the prestressed FRP U-straps to increase their
effectiveness. This may be a subject for further research. Although unconfirmed by
further experimental studies, the slight advantages observed from prestressing are
outweighed by their labor intensiveness and poor practicality.
2.3.4 Metallic Anchorage Systems
Metallic anchorages are one of the earliest forms of FRP end anchorage devices
investigated by researchers (e.g. Sharif et al. 1994; Jensen et al. 1999). Investigations
have been conducted on adhesively bonded metallic plates with mechanical fasteners
(Figure 2.4), adhesively bonded metallic U-jackets, and U-jackets with end clamping.
Researchers such as Garden and Hollaway (1998), Spadea et al. (1998), Duthinh and
Chapter 2 – Literature Review of FRP Anchorage Systems in concrete Infrastructure
15
Starnes (2001), and Wu and Huang (2008) have found that the use of metallic
anchorages provides a significant increase in anchorage strength in addition to ductility
enhancement.
Figure 2.4 - (a) Typical FRP plate anchored using permanent mechanical anchorage device (b) schematic of typical test setup
Previous experimental testing demonstrated the ineffectiveness of bonded angle sections
for plate-end anchorage due to the lack of a secure plate end fixing to the concrete.
Experiments were conducted by Garden and Hollaway (1998) with a number of 1.0 m
long plated beams tested in four-point bending. Cantilevers were also tested to
demonstrate that the structural benefit of plate-end anchorage diminishes as the shear
span/depth ratio of the beam increases. Each beam was strengthened with 67 mm wide
and 0.87 mm thick, 111–115 GPa modulus CFRP plates. The bolted plate-end
anchorage system used comprised a 40 mm long steel anchorage block of the same
width as the composite plate. The block was secured to the composite plate using
laminate adhesive and two mild steel bolts.
A comparison was made between the mechanically fastened steel anchorages and where
the bonded plate was continued under the supports of the beam, resulting in a clamping
force applied normal to the plate. The authors concluded that the main requirements of
bolted plate-end anchors were the shear resistance of the anchor bolts and the FRP-steel
adhesive bond. The conclusion was based upon the similarity of the results obtained
between clamping and fastening anchors. The authors did not compare fastened steel
anchors with unclamped, unfastened anchors, which would be needed to prove that
Chapter 2 – Literature Review of FRP Anchorage Systems in concrete Infrastructure
16
confinement does not improve anchorage effectiveness. Because the combined benefits
of bolted plates together with clamping pressure were not investigated, the benefits of
the application of clamping forces together with mechanical fastening remain to be fully
substantiated. Duthinh and Starnes (2001) tested a series of seven beams in four-point
bending. A single carbon fiber plate (1.2 × 50 mm ) with an elastic modulus of 155 GPa
was used to strengthen the beams in flexure. Three of the beams tested utilized a 203
mm wide mechanically fastened steel anchor over the plate end. Two bolts were torqued
to 400 Nm, resulting in an applied clamping force of 15–25 kN. The result of clamping
and adhesion enabled the anchored plate to reach an ultimate strain of 11400 (60% of
rupture). Failure was by debonding initiating from diagonal shear cracking. The authors
stipulated that clamping combined with adhesion can double or triple the anchorage
capacity that can be expected from the bond alone. However, no investigations were
carried out using bolted anchorages without torque to assess the contribution of
clamping force on anchorage enhancement within the context of the test setup.
Spadea et al. (1998) attempted to improve the performance of CFRP-strengthened RC
beams by using external steel anchorages designed to control and minimize the bond-
slip between the concrete beam and the CFRP plate. The anchorages consisted of U-
shaped steel anchors installed at the plate ends, together with four to eight U-shaped
steel anchorages distributed throughout the span. The plates were bonded to the
concrete using epoxy resin and contained no external bolts or mechanical fasteners.
Experimental testing measured maximum fiber strain utilizations of 80% (12,000 )
for beam specimens with end anchorages at the plate ends, together with eight U-
shaped anchorages distributed throughout the span, corresponding to a 67%
enhancement over the corresponding unanchored specimen. In addition to the enhanced
fiber utilization and strength enhancement provided by the steel anchorages, greater
ductility and gradual debonding of the plate over an extended time increment were also
observed. Ductility was evaluated through an examination of deflection (deflection
ductility), curvature (curvature ductility), and the area-under-the-load deflection
curve at yielding of the tension steel and ultimate failure (energy ductility). The
detailing of bonded CFRP plates without anchorage was found to reduce the ductility
index by 70–80%, whereas when provisions were made for ad- equate anchorage, the
loss of ductility was only 45–70%. Although the improvements in ductility are very
Chapter 2 – Literature Review of FRP Anchorage Systems in concrete Infrastructure
17
attractive to designers, the wide range of ductility indices indicates that a more
consistent approach is required to define and quantify the ductility of FRP-
strengthened beams. The strain in the tensile reinforcement at failure was not measured.
Researchers have attempted to combine the benefits of mechanically fastened (MF-
FRP) systems with the traditional externally bonded (EB-FRP) method, resulting in a
new hybrid plate (HB-FRP) bonding system (Wu and Huang 2008). The fasteners used
in this study are presented in figure 2.5.
Figure 2.5 - (a) mechanical fastener; (b) predrilled holes; (c) Details of the HB-FRP system; adapted from (Wu and Huang 2008).
The application of the HB-FRP system comprises initially the attachment of the FRP to
the concrete surface using adhesive after adequate surface preparation. Following full
curing of the adhesive, special mechanical fasteners are installed longitudinally along
the FRP reinforcement at a specified spacing. Insertion of the mechanical fasteners
follows the same procedure as the MF-FRP method. The fasteners do not carry any
bearing forces, but act to increase the bond strength between the FRP and the concrete
by resisting the tensile peeling stresses which can initiate a debonding failure.
Wu and Huang (2008) observed two distinct failure modes of the hybrid system, namely
(1) CFRP rupture at midspan, which occurred with specimens strengthened with 2- and
4-ply strips, and (2) complete strip debonding, which was observed for the specimen
strengthened with 6-ply strips, indicating that the bond strength had been exhausted.
Considerable increases in flexural capacity and bond strength were observed as a result
of the hybrid plate-bonding system. A 79% increase in moment resistance was
attributed to the addition of the fasteners alone for the same area of CFRP. However, the
increase in bond strength was even higher than the moment increase. This resulted in
specimens mechanically fastened with 4-ply and 6-ply strips reaching flexural strengths
of 184.9% and 268.2%, respectively, higher than the 2-ply specimen with no fasteners.
Chapter 2 – Literature Review of FRP Anchorage Systems in concrete Infrastructure
18
The application of steel anchorages to CFRP strengthened members is limited by factors
such as cost, practicality, labor intensiveness, and durability. Drilling threaded rods or
expansion anchors into existing structures is time-consuming and has the potential to
damage existing reinforcement. In addition, long-term durability is a concern and is
aggravated by the galvanic coupling with the carbon fiber, which must be mitigated by
use of a glass fiber layer between the steel and the concrete. Research has demonstrated
that steel anchorages generally provide higher anchorage strength than non metallic
anchors because of their metallic rigidity and the ability of mechanical fasters to
effectively resist tensile and shear forces.
2.3.5 FRP Anchors
Anchors made from rolled fiber sheets or bundled loose fibers are a promising form of
anchorage because they can be applied to wide shaped FRP strengthened structural
elements such as slabs and walls. They are discrete and do not suffer from the same
constraints as U-jackets. Such anchors are referred to as FRP spike anchors, fiber
anchors, fiber bolts and FRP dowels, amongst other names, but are herein collectively
referred to FRP anchors (Smith 2010). The anchor can be hand-made (in the laboratory
or on site) or manufactured from glass or carbon fiber sheets or loose fibers which have
been rolled or bundled. Such method of manufacturing makes the anchors extremely
simple to construct but quite variable (especially if hand-made). The variation, however,
does not appreciably affect the behaviour of the anchored EB-FRP system (Zhang et al.
2010). As indicated in Figure 2.6a, one end of the anchor (herein anchor dowel) is
inserted into a pre-drilled hole in the concrete substrate and the dowel length can be
confined to the cover region of the member. The other end of the anchor is epoxied onto
the surface of the EB-FRP. The ends of the fibers which are splayed and epoxied onto
the surface of the plate in order to disperse local stress concentrations are herein referred
to as the anchor fan.
A convenient means by which to determine the fundamental strength and behavioural
characteristics of FRP anchors is to test them in FRP-to-concrete joint assemblies such
as that shown in Figure 2.6(d), from Zhang et al. (2012) and several researchers have
Chapter 2 – Literature Review of FRP Anchorage Systems in concrete Infrastructure
19
investigated such joints to date (e.g. Zhang et al. 2012; Zhang and Smith 2012a, b;
Niemitz 2008). A generic load-slip response of single fan and bow-tie anchors is shown
in Figure 2.6(e). The three main stages of the load-slip response are denoted by A (i.e.,
debonding and activation of FRP anchor), B (i.e., post peak reserve of strength offered
by completely intact FRP anchor and frictional resistance of debonded plate), and C
(i.e., post peak reserve of strength offered by partially intact FRP anchor and frictional
resistance of debonded plate). Ongoing research is establishing the key loads (P) and
slips ( ) for varying anchor material and geometric properties (e.g., Kim and Smith
2009; Smith 2010; Zhang et al. 2012). A review by Smith (2010) reported that FRP
spike anchors with a single fan component increase the shear strength and slip capacity
of FRP-to-concrete joints by up to 70% and 800%, respectively, over unanchored
control joints. Of particular interest in Figure 2.6(f) is the significant effect of dowel
angle on the joint strength enhancement over the unanchored control joint (Zhang and
Smith 2012a). One of the earliest reported tests on FRP anchors in a concrete member
was by Lam and Teng (2001). In their work, RC cantilever slabs of 700 mm span
strengthened with glass FRP (GFRP) plate bonded to the tension face of the slabs were
tested. The use of a GFRP anchor as a mechanical anchorage system can also prevent
premature peeling of CFRP laminates in the presence of curvature.
Chapter 2 – Literature Review of FRP Anchorage Systems in concrete Infrastructure
20
Figure 2.6 - (a, b, c) Anchor construction and installation of FRP anchors (reprinted from Engineering Structures, Vol. 33, No. 4, Smith, ST, Hu, S, Kim, SJ & Seracino, R 2011, “FRP-strengthened Rc slabs anchored with FRP anchors”, Pages 1075–1087, April 2011, with permission from Elsevier); (d) test setup (single lap) (reprinted from Construction and Building Materials, FRPRCS9 Special Edition, H.W. Zhang, S.T. Smith, S.J. Kim, “Optimisation of carbon and glass FRP anchor design”, Pages 1–12, June 2012, with permission from Elsevier); (e) generic load-slip response of FRP-to-concrete joint anchored with bow-tie anchor; (f) joint strength enhancement (above unanchored control) [modified from Zhang and Smith (2012b)] Eshwar et al. (2005) investigated 200 × 400 mm RC beams spanning 5.5 m with both
straight and curved beam soffits (curvature 5 mm over 1 m). A single row of 10 mm
FRP spike anchors was embedded 76 mm into the concrete beam at 500 mm spacing’s.
Reductions in strength of 20% and 30% were observed in beams strengthened with wet
lay-up fibers and precured laminate due to curvature and premature peeling. Inclusion
of the anchor FRPs with the wet lay-up system applied to the curved-soffit specimen led
to the strength being increased by 35% compared to the unanchored specimen. This
resulted in the strength of the curved-soffit beam containing the anchor FRPs being
higher than that of the flat soffit beam strengthened with wet lay-up fibers. Others have
investigated the performance of FRP anchors in flexural members (e.g., Micelli et al.
2010). In most cases, the addition of FRP anchors was found to increase the strength
and ductility of the FRP-strengthened members. However, this is not always the case
and reasons why remain to be addressed.
Chapter 2 – Literature Review of FRP Anchorage Systems in concrete Infrastructure
21
Further research has shown that the use of FRP anchors is an effective way to improve
the strength of reinforced concrete members. Orton et al. (2008) determined that two
rows of three 10 mm diameter anchors were able to develop the FRP tensile capacity
and led to fracture of the entire width of the FRP. They reported that FRP anchors
increased the efficiency of material usage of the FRP retrofit by 57%, indicating that
FRPs with anchors are able to achieve a given strengthening capacity and require less
material than unanchored FRPs. In this case, the strength of the member increased by
270%, with only a 175% increase in the FRP material. In addition, it was found that a
greater number of smaller anchors and reduced spacings were more effective in fully
developing the capacity of the FRP fiber, as larger spacings did not anchor the entire
width of the FRPs, resulting in partial debonding (Orton et al. 2008).
Lam and Teng (2001) conducted investigations on improving the strength of wall
cantilever slab connections using GFRP strips. Fiber anchors were installed to anchor
the GFRP strips into the RC wall. The authors observed that debonding was stopped by
the fiber anchors and the slabs finally failed by tensile rupture of the FRP. In tests on
similar slabs simply bonded with two 80.5-mm wide GFRP strips without the use of
fiber anchors, debonding between the FRP and the slab occurred in all cases (Teng et al.
2000).
2.3.6 Evaluation of FRP anchors used to strengthen members in flexure.
(Grelle and Sneed 2011) recently established the need to establish a large database of
anchorage test results. This section therefore presents a database of selected strain data
for FRP anchorage systems, where each anchorage type can be compared using a
common correlation parameter. In order to comparatively assess each anchorage, the
concrete strength (f’c), fiber modulus (Ef), number of plies (n) and fiber thickness (tf),
were used to standardise the strain data from experimental results collected from a
number of researchers which is presented in table 2.1. Fiber modulus, number of plies
and fiber thickness all affect the magnitude of FRP-to-concrete bond stresses at the
interface at a given level of FRP strain, whereas the concrete strength is the key
parameter which governs the bond resistance of the interface. It is therefore important to
consider these factors when determining the strain efficiency of any strengthened
Chapter 2 – Literature Review of FRP Anchorage Systems in concrete Infrastructure
22
system. An anchorage effectiveness factor has been defined on the basis of the
maximum strain reached in the FRP plate prior to failure, f,max, and the effective FRP
strain to resist intermediate crack debond, f,d (ACI 440.2R-08 2008). The resulting
expression presented in equation 2.1, which is used to define the anchorage
effectiveness factor (kfab).
(2.1)
Comparing anchorages in this manner can provide a concise behavioural summary of
alternative anchorage solutions with respect to FRP strain efficiency. Factors such as the
limited number of test specimens for the majority of experimental regimes weaken the
statistical reliability of the database. This shortcoming can only be addressed once more
data becomes available. However, the results may still serve as a useful comparison of
available anchorage methods. In addition, equation 2.1 does not take into account
mechanical parameters not included in the equation, as well as the quality of
workmanship in preparing the specimens. As a result of reviewing various experimental
procedures and results currently published, it was found that in many instances the data
was not utilized due to specimens failing either by concrete crushing, or a failure to
present or measure the strain in the FRP prior to failure and the corresponding strain in
the FRP anchorages. This strain data is especially useful when assessing anchorage
behaviour. It is suggested that all future research in this area make use of under rein-
forced sections for flexurally strengthened members to ensure that specimen failure
occurs by either FRP debonding or FRP rupture and presents adequate FRP strain
measurement data for use by other researchers.
Of the various anchorage types listed to improve the flexural efficiency of FRP-
strengthened beams, metallic anchorages are found to be the most effective, in which
maximum fiber elongation reached prior to failure is the sole evaluation criteria.
Inclined U-jacket anchors, are observed to be 65% more effective than the traditional U-
jacket anchors, resulting in exceptional anchorage efficiency kfab = 2.42. U-jackets are
attractive due to their simplicity, non-destructiveness, and ease of installation, making
them ideal choices for T-beam applications. When comparing pre- stressed FRP U-
jackets within the context of the Pham and Al-Mahaidi (2006) program, the
Chapter 2 – Literature Review of FRP Anchorage Systems in concrete Infrastructure
23
anchorages failed prematurely due to lack of adequate restraint of the U-strap ends As a
result, the relatively low kfab factor observed may not be representative of the full
potential of prestressing. In principle, it is expected that prestressed U-straps should
always result in higher anchorage efficiency due to the higher degree of confinement
and shear resistance provided within the anchorage zone. This result is expected to be
improved upon once a more effective anchorage arrangement is provided to the ends of
the U-straps, a subject of further research.
FRP anchors were found to be third highest in efficiency based on limited test data (kfab
= 2.03) and have also been shown to significantly enhance deformability and ductility.
The slip capacity of such joints has also been observed to increase by several hundred
percent. FRP anchors have the highest flexibility and potential for application to both
slab and beam members, and their effectiveness and ease of installation make them a
highly recommended form of anchorage.
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k f
a Fa
ilure
1
M
Pa
mm
G
Pa
FR
P Fl
exur
al fi
ber
only
0.
58
(Ave
rage
) (P
iman
mas
200
3)
A-2
00P
200m
m S
uppo
rt 55
.0
1.20
15
0.0
3860
0.
54
IC
(Pim
anm
as 2
003)
A
-200
P 42
0mm
Sup
port
55.0
1.
20
150.
0 34
20
0.48
ED
(P
iman
mas
200
3)
B-20
0P
200m
m S
uppo
rt 55
.0
1.20
15
0.0
2890
0.
40
ED
(Pha
m a
nd A
l-Mah
aidi
200
6)
E1a
6 PL
Y -
3 x
12m
m d
ia b
ars
53.7
1.
06
209.
0 30
36
0.47
ED
(P
ham
and
Al-M
ahai
di 2
006)
E3
a 6
PLY
- 2
x 12
mm
dia
bar
s 53
.7
1.06
20
9.0
3502
0.
55
ED
(Pha
m a
nd A
l-Mah
aidi
200
6)
E1b
6 - 3
x 1
2mm
dia
bar
s 53
.7
1.06
20
9.0
3414
0.
53
ED
(Pha
m a
nd A
l-Mah
aidi
200
6)
E5a
9 PL
Y C
FRP
53.7
1.
06
209.
0 23
29
0.36
ED
(S
mith
et a
l. 20
11)
S2
Un-
anch
ored
con
trol
41.4
0.
50
239.
0 66
49
0.87
IC
(Y
alim
, Kal
ayci
et a
l. 20
08)
W1.
1 C
FRP
– su
rfac
e sm
ooth
(CS1
) 35
.0
1.02
70
.5
6039
0.
67
IC
(Yal
im, K
alay
ci e
t al.
2008
) W
1.2
Surf
ace
(CS1
) 35
.0
1.02
70
.5
7443
0.
82
IC
(Yal
im, K
alay
ci e
t al.
2008
) W
2.3.
1 Su
rfac
e (C
S2-C
S3)
35.0
1.
02
70.5
64
90
0.72
IC
(Y
alim
, Kal
ayci
et a
l. 20
08)
W6.
9.1
Surf
ace
(CS6
-CS9
) 35
.0
1.02
70
.5
5214
0.
58
IC
FRP
U-j
acke
t Anc
hor
0.78
(A
vera
ge)
(Yal
im, K
alay
ci e
t al.
2008
) P1
.1
4 C
FRP
U-ja
cket
s 35
.0
1.40
13
1.0
4842
0.
85
IC
(Yal
im, K
alay
ci e
t al.
2008
) P2
.3.1
4
CFR
P U
-jack
ets
35.0
1.
40
131.
0 45
98
0.81
IC
(Y
alim
, Kal
ayci
et a
l. 20
08)
P6.9
.1
4 C
FRP
U-ja
cket
s 35
.0
1.40
13
1.0
5027
0.
89
IC
(Yal
im, K
alay
ci e
t al.
2008
) P2
.3.2
Fu
ll U
-jack
et
35.0
1.
40
131.
0 50
76
0.90
IC
(Y
alim
, Kal
ayci
et a
l. 20
08)
P6.9
.2
Full
U-ja
cket
35
.0
1.40
13
1.0
5281
0.
93
IC
(Pim
anm
as 2
003)
A
-420
U
90 d
egre
e U
-jack
et a
ncho
r 55
.0
1.20
15
0.0
8760
1.
22
CC
/ ED
(P
iman
mas
200
3)
B-2
00U
90
deg
ree
U-ja
cket
anc
hor
55.0
1.
20
150.
0 37
50
0.52
C
C /
IC
(Pha
m a
nd A
l-Mah
aidi
200
6)
A1a
1
U-ja
cket
- 3
x 12
mm
dia
bar
s 53
.7
1.06
20
9.0
4100
0.
64
IC
(Pha
m a
nd A
l-Mah
aidi
200
6)
A1b
3
U-ja
cket
s at 1
80 m
m c
/c -
3 x
12m
m d
ia b
ars
53.7
1.
06
209.
0 53
50
0.84
IC
(P
ham
and
Al-M
ahai
di 2
006)
E3
a2
1 U
-jack
et -
2 x
12m
m d
ia b
ars
53.7
1.
06
209.
0 35
00
0.55
IC
(P
ham
and
Al-M
ahai
di 2
006)
E5
a2
3 U
-jack
ets a
t 180
mm
c/c
- 3
x 12
mm
dia
bar
s 53
.7
1.69
20
9.0
4307
0.
83
IC
(Yal
im, K
alay
ci e
t al.
2008
) W
1.3
4 U
-jack
ets 2
No.
EA
CH E
ND
. (C
S1)
35.0
1.
02
70.5
63
14
0.70
ED
(Y
alim
, Kal
ayci
et a
l. 20
08)
W1.
4 4
U-ja
cket
s 2 N
o. E
ACH
EN
D. (
CS1
) 35
.0
1.02
70
.5
3876
0.
43
ED
(Yal
im, K
alay
ci e
t al.
2008
) W
1.5
4 U
-jack
ets 2
No.
EA
CH E
ND
. (C
S1)
35.0
1.
02
70.5
66
85
0.74
ED
(Y
alim
, Kal
ayci
et a
l. 20
08)
W2.
3.2
4 C
FRP
U-ja
cket
s 2 N
o. E
AC
H E
ND
. (CS
2-C
S3)
35.0
1.
02
70.5
77
91
0.86
ED
(Y
alim
, Kal
ayci
et a
l. 20
08)
W2.
3.3
4 C
FRP
U-ja
cket
s 2 N
o. E
AC
H E
ND
. (CS
2-C
S3)
35.0
1.
02
70.5
73
86
0.82
ED
(Y
alim
, Kal
ayci
et a
l. 20
08)
W2.
3.4
4 C
FRP
U-ja
cket
s 2 N
o. E
AC
H E
ND
. (CS
2-C
S3)
35.0
1.
02
70.5
68
14
0.75
ED
(Y
alim
, Kal
ayci
et a
l. 20
08)
W6.
9.2
4 C
FRP
U-ja
cket
s 2 N
o. E
AC
H E
ND
. (CS
6-C
S9)
35.0
1.
02
70.5
80
57
0.89
ED
(Y
alim
, Kal
ayci
et a
l. 20
08)
W6.
9.3
4 C
FRP
U-ja
cket
s 2 N
o. E
AC
H E
ND
. (CS
6-C
S9)
35.0
1.
02
70.5
62
53
0.69
ED
(Y
alim
, Kal
ayci
et a
l. 20
08)
W6.
9.4
4 C
FRP
U-ja
cket
s 2 N
o. E
AC
H E
ND
. (CS
6-C
S9)
35.0
1.
02
70.5
64
22
0.71
ED
(Y
alim
, Kal
ayci
et a
l. 20
08)
W1.
6 7
CFR
P U
-jack
ets (
CS1
) 35
.0
1.02
70
.5
8349
0.
92
ED
(Yal
im, K
alay
ci e
t al.
2008
) W
1.7
11 C
FRP
U-ja
cket
s (C
S1)
35.0
1.
02
70.5
89
62
0.99
FR
(Y
alim
, Kal
ayci
et a
l. 20
08)
W2.
3.5
11 C
FRP
U-ja
cket
s (C
S2-C
S3)
35.0
1.
02
70.5
83
81
0.93
FR
(Y
alim
, Kal
ayci
et a
l. 20
08)
W6.
9.5
11 C
FRP
U-ja
cket
s (C
S6-C
S9)
35.0
1.
02
70.5
10
074
1.11
FR
(Y
alim
, Kal
ayci
et a
l. 20
08)
W1.
8 Fl
exur
al F
RP
+ Fu
ll U
-jack
et (C
S1)
35.0
1.
02
70.5
66
47
0.73
FR
(Y
alim
, Kal
ayci
et a
l. 20
08)
W2.
3.6
Full
U-ja
cket
(CS2
-CS3
) 35
.0
1.02
70
.5
8937
0.
99
FR
1 CC
= C
oncr
ete
Cru
shin
g; IC
= In
term
edia
te C
rack
Indu
ced
Deb
ondi
ng; F
R =
Fib
er R
uptu
re; E
D =
End
Deb
ond;
; ES
= En
d Sl
ippa
ge ;
Cha
pter
2 –
Lite
ratu
re R
evie
w o
f FRP
Anc
hora
ge S
yste
ms i
n co
ncre
te In
fras
truc
ture
25
Tab
le 2
.1 -
FRP
anch
orag
e su
mm
ary
for f
lexu
rally
stre
ngth
ened
mem
bers
Sp
ecim
en
Com
men
ts
f' c
t ft
E f
f,max
k f
a Fa
ilure
1
M
Pa
mm
G
Pa
FR
P Fl
exur
al fi
ber
only
0.
58
(Ave
rage
) (P
an, L
eung
et a
l. 20
10)
B1
Sing
le n
otch
ed b
eam
with
side
pla
tes
49.2
0.
22
235.
0 66
28
0.52
IC
(P
an, L
eung
et a
l. 20
10)
B2
Sing
le n
otch
ed b
eam
with
side
pla
tes
49.2
0.
22
235.
0 66
25
0.52
IC
(P
an, L
eung
et a
l. 20
10)
B3
Dou
ble
notc
hed
beam
with
side
pla
tes
49.2
0.
22
235.
0 72
99
0.58
IC
(P
an, L
eung
et a
l. 20
10)
B4
Dou
ble
notc
hed
beam
with
side
pla
tes
49.2
0.
22
235.
0 64
92
0.51
IC
(P
an, L
eung
et a
l. 20
10)
B5
Dou
ble
notc
hed
beam
with
FR
P pl
ate
49.2
0.
22
235.
0 10
217
0.81
IC
(P
an, L
eung
et a
l. 20
10)
B6
Un-
notc
hed
beam
with
FR
P pl
ate
49.2
0.
22
235.
0 10
489
0.83
IC
(P
an, L
eung
et a
l. 20
10)
B7
Pre-
crac
ked
bond
ed w
ith F
RP
plat
e 49
.2
0.22
23
5.0
9399
0.
74
IC
(Pan
, Leu
ng e
t al.
2010
) B
8 U
n-no
tche
d be
am w
ith F
RP
plat
e 49
.2
0.22
23
5.0
9954
0.
79
IC
Pres
tres
sed
U-j
acke
t Anc
hor
0.78
(A
vera
ge)
(Pha
m a
nd A
l-Mah
aidi
200
6)
A2a
1
Pres
tress
ed U
-jack
et -
3 x
12m
m d
ia b
ars
53.7
1.
06
209.
0 45
71
0.71
IC
(P
ham
and
Al-M
ahai
di 2
006)
A
2b
3 Pr
estre
ssed
U-ja
cket
s at
180
mm
c/c
- 3
x
53.7
1.
06
209.
0 54
16
0.85
IC
In
clin
ed F
RP
U-j
acke
t Anc
hor
1.36
(A
vera
ge)
(Sag
awa,
Mat
sush
ita e
t al.
2001
) 200
1)
U1-
45-1
In
clin
ed U
-jack
et a
ncho
r, 1
plac
e 27
.3
0.17
23
0.0
1500
0 1.
36
FR
(Sag
awa,
Mat
sush
ita e
t al.
2001
) U
1-45
-2
Incl
ined
U-ja
cket
anc
hor,
2 pl
aces
27
.3
0.17
23
0.0
1500
0 1.
36
FR
FRP
+ St
eel A
ncho
rage
1.
87
(Ave
rage
) (S
pade
a, B
enca
rdin
o et
al.
2000
) A
1.2
Stee
l Anc
hora
ges T
ype
A/T
ype
B
30.0
1.
20
152.
0 96
00
1.83
E
D
(Spa
dea,
Ben
card
ino
et a
l. 20
00)
A1.
3 St
eel A
ncho
rage
s Typ
e A
/Typ
e B
/Typ
e C
30
.0
1.20
15
2.0
1050
0 2.
00
ES /
ED
(Spa
dea,
Ben
card
ino
et a
l. 20
00)
A2.
2 St
eel A
ncho
rage
s Typ
e A
/Typ
e B
- A
rr1
30.0
1.
20
152.
0 10
000
1.90
ES
/ ED
(S
pade
a, B
enca
rdin
o et
al.
2000
) A
2.3
Stee
l Anc
hora
ges T
ype
A/T
ype
B -
Arr
2 30
.0
1.20
15
2.0
1100
0 2.
09
ES/E
D/C
C
(Spa
dea,
Ben
card
ino
et a
l. 20
00)
A3.
2 St
eel A
ncho
rage
s Typ
e A
/Typ
e B
30
.0
1.20
15
2.0
1020
0 1.
94
ED
(S
pade
a, B
enca
rdin
o et
al.
2000
) A
3.3
Stee
l Anc
hora
ges T
ype
A/T
ype
B/T
ype
C
30.0
1.
20
152.
0 12
000
2.28
ES
(D
uthi
nh a
nd S
tarn
es 2
001)
B
4a
Stee
l Cla
mp
at L
amin
ate
ends
, 400
N.m
42
.3
1.20
15
5.0
1007
0 1.
63
ED
(Dut
hinh
and
Sta
rnes
200
1)
B6
Stee
l Cla
mp
at L
amin
ate
ends
, 400
N.m
41
.3
1.20
15
5.0
7800
1.
28
ES
FRP
Anc
hors
1.
14
(Ave
rage
) Sm
ith e
t al.
(201
1)
S3
FRP
anch
ors a
long
who
le sp
an (T
ype
A)
41.4
0.
50
239.
0 76
76
1.00
IC
Sm
ith e
t al.
(201
1)
S4
FRP
anch
ors a
long
who
le sp
an (h
alf n
o. a
ncho
r as
44.1
0.
50
239.
0 80
25
1.02
IC
Sm
ith e
t al.
(201
1)
S5
Shea
r spa
n FR
P an
chor
s (Ty
pe A
) 44
.1
0.50
23
9.0
8884
1.
13
IC
Smith
et a
l. (2
011)
S6
Pl
ate
end
FRP
anch
ors (
Type
A)
45.4
0.
50
239.
0 66
96
0.84
IC
Sm
ith e
t al.
(201
1)
S7
Shea
r spa
n FR
P an
chor
s (Ty
pe B
) 45
.4
0.50
23
9.0
1156
6 1.
44
IC
Smith
et a
l. (2
011)
S8
Sh
ear s
pan
FRP
anch
ors (
Type
A +
Typ
e B
) 45
.4
0.50
23
9.0
1134
8 1.
42
IC
1 CC
= C
oncr
ete
Cru
shin
g; IC
= In
term
edia
te C
rack
Indu
ced
Deb
ondi
ng; F
R =
Fib
er R
uptu
re; E
D =
End
Deb
ond;
; ES
= En
d Sl
ippa
ge ;
Chapter 2 – Literature Review of FRP Anchorage Systems in concrete Infrastructure
26
2.4 Mechanisms of FRP failure in shear retrofit applications
Common techniques for strengthening RC members in shear using FRP are: side
bonding, U-jacketing and full wrapping. Experience has shown that failure of FRP
bonded to concrete as externally bonded shear reinforcement is closely related to the
shear strengthening system utilised. The majority of experimental data highlights that
almost all beams strengthened by enclosed wrapping typically fail due to FRP rupture
after localised debonding (Chen and Teng 2003). In contrast, beams strengthened by
side bonding only and most strengthened by U jacketing, fail due to debonding of the
FRP, which has been observed to initiate where the FRP intersects diagonal shear
cracks in the member. Debonding then propagates to the nearer end of the plate (this is
typically the free plate end). It may be noted that pure interfacial debonding failure
along the FRP-adhesive interface, adhesive-concrete interface or within the adhesive
have been rarely reported. Debonding failures almost always occur within the concrete
at the FRP-to-concrete interface.
2.5 Anchorage devices for FRP reinforcement used to strengthen members in shear
Although fully wrapping the beam cross-section with FRP has been demonstrated to
provide the most effective strengthening solution for shear and torsion applications, it is
seldom achieved in practice due to the presence of physical obstructions such as beam
flanges. U-jacketing is currently the most popular shear strengthening solution because
of its high practicality, but it is limited by end-peeling of the U-jacket legs. This
form of failure is usually premature, sudden, and non-ductile, and it has resulted in the
development of many innovative anchorage details at the web- flange interface. These
include the following:
1. FRP enveloping the web of the beam in a U-shape, including termination at the
underside of the beam flange with no anchorage (Khalifa et al. 2000; Micelli et al. 2002;
Tanarslan et al. 2008).
2. Wrapping the web and flange of the beams through drilled holes through the beam
flanges (Hoult and Lees 2009).
Chapter 2 – Literature Review of FRP Anchorage Systems in concrete Infrastructure
27
3. Mechanically fastened metallic anchors installed at the under- side of the beam
flange to anchor FRP U-wrap legs (Deifalla and Ghobarah 2010; Micelli et al. 2002;
Tanarslan et al. 2008).
4. Embedment of the FRP U-jacket legs into the beam flanges through pre-cut grooves
using adhesive bonding (Lee and Al-Mahaidi 2008).
5. FRP anchors installed to restrain the legs of the FRP U-jackets.
2.5.1 Mechanically fastened metallic anchors in shear and torsion applications
The efficiency of metallic anchorages has been found to be case-dependent and less
suitable in shear and torsion retrofits. The subject was investigated by Panchacharam
and Belarbi (2002), who tested eight beams in pure torsion. The strengthening schemes
included: complete wrapping, U-jacketing, and U-jacketing with mechanically fastened
metallic anchors. The inefficiency of U-jackets applied to rectangular beams subjected
to torsion was verified by the 80% increase in torsional resistance when complete
wrapping was provided compared to that of U-jackets only. The author reported no
increase in ultimate strength between U-jacketed test beams strengthened with and
without mechanical anchorages. The presence of anchors was, however, found to
increase the post cracking twist and energy absorption capacity when compared to
unanchored U-jacketed test beams. The results suggest that in torsion applications, FRP
U-jackets are a poor alternative to full wrapping, even when mechanical anchorage is
provided.
Similar research conducted on concrete T-beams loaded in pure torsion has verified the
ineffectiveness of metallic anchors to improve the performance of FRP U-jacket strain
levels (Salom et al. 2004). However, a higher torsion capacity was achieved due to the
fastening of the metallic anchorage to the underside of the T-beam flanges. This was
attributed to the anchor bolts acting as a part of the shear flow mechanism and was
verified by the high strain values recorded in the anchor bolts.
Deifalla and Ghobarah (2010) evaluated a mechanically anchored extended U-jacket
system by investigating six concrete T-beams subjected to combined shear and torsion
Chapter 2 – Literature Review of FRP Anchorage Systems in concrete Infrastructure
28
in a configuration similar to that shown in Figure 2.7. The experiments utilized a
bidirectional carbon composite fiber with ±45° fiber orientation and a modulus of
elasticity of 63.3 GPa. In this technique, the U-jacket was bonded to the web of the
beam and anchored 50 mm below the intersection of the web and the flange. An
additional steel angle fastened to the beam flange with 20 mm diameter steel threaded
rods was used at the en- trance of the flange and the web to delay end-jacket debonding
failure. Using the extended U-jacket together with mechanically fastened steel angles
was found to be more effective than using the U-jacket anchored to the beam web with
20 mm rods only.
Figure 2.7 - Implemented strengthening schemes (a) U-jacket; (b) Extended U-jacket; adapted from (Deifalla and Ghobarah 2010)
A 23% increase in strength and an enhanced ductility of 38% were achieved compared
to that of the web-anchored U-jacket technique. Ductility was measured by considering
both deflection and twist ductility (monitoring the maximum angle of twist) and the
maximum strain level of the steel reinforcement. The authors suggested that the
enhanced torsion capacity was because of an increase of the enclosed area inside the
expected critical shear flow path induced by the mechanical anchorage provided into
the beam flanges. However, no comparisons with unanchored U-jacketed specimens
were made to assess the contributions of the steel anchorages.
Mechanically anchored U-jackets have achieved greater effectiveness in pure
shear applications (Aridome et al. 1998; Maeda et al. 1997; Ortega et al. 2009;
Tanarslan et al. 2008). An investigation into the shear behaviour of concrete T-beams
strengthened with alternative CFRP schemes was conducted by Tanarslan et al. (2008).
Chapter 2 – Literature Review of FRP Anchorage Systems in concrete Infrastructure
29
The study encompassed specimens retrofitted with CFRP side bonding, L-wrapping (leg
of L developed beneath flange), U-jacketing, and extended U-jacketing. Steel
anchorages were applied to CFRP sheets in both top and bottom locations for four of the
specimens tested. In addition, 10 mm threaded rods were used to fasten the 50 × 50 × 5
mm steel plates at CFRP soffit terminations and L-shaped 50 × 50 × 5 mm steel plates
were used at the web/flange interfaces. L-shaped strips with anchorage prevented
premature debonding
but failed prematurely due to tearing of the concrete cover below the level of the bottom
reinforcement. This mode of failure indicates that a development of side-bonded FRP
below the beam soffit is required for anchorages to achieve their full potential. The
failure mode was prevented in the anchored U-jacketed specimens, which achieved an
additional 35% in shear capacity over L-wrapping and failed through shear crack-
induced FRP rupture. Although the anchored extended U-jacket showed the highest re-
corded shear strength, the increased FRP width used for the specimen makes
comparative observations difficult. It is recommended that future research should
always utilize consistent FRP material properties and dimensions to enable accurate
correlations to be made between alternative anchorage techniques in any given program.
The effect of using continuous and discontinuous steel/CFRP plates bonded to the top
and bottom of shear reinforcement was investigated by Ortega et al. (2009). The
steel/CFRP plate anchors were fixed using concrete wedge anchors and steel bolts. A
typical representation is shown in Figure 2.8. In this study, continuous mechanically
fastened steel plate anchorages were ineffective because the continuous plate exhibited
a bucking failure mode due to the curvature of the beam at failure. The fasteners
exhibited bearing failure in some locations. In addition, slippage of the CFRP prevented
the CFRP shear reinforcement from reaching its full capacity. This was solved by the
development of a modified anchor bolt system, which consisted of wrapping the CFRP
composite around the first plate and overlapping with the second plate, creating a three-
layer connection.
Chapter 2 – Literature Review of FRP Anchorage Systems in concrete Infrastructure
30
Figure 2.8 - View Anchorage System with discontinuous steel anchorages, adapted from (Ortega et al. 2009).
This behaviour was also verified by Aridome et al. (1998), who concluded that
continuous steel plate anchors separated prematurely due to in-plane bending stresses
within the steel anchorage. Staggered plate anchors were found to provide the highest
beam ductility, which was measured by monitoring beam deflections. To equate vertical
deflections with ductility is not representative of the beam’s ability to undergo sufficient
cracking and deformability prior to failure. Cracking and deformability are the current
measures used to ensure ductility in FRP-strengthened members in FRP design
guidelines monitored by the strain level in the tensile reinforcement. The staggering of
steel anchorages within the compression zone was important to reduce the overall
compression block stiffness, resulting in higher deflections. However, as a result of
plate staggering, the compression block stiffness shifts the neutral axis of the section
toward the bottom fiber, resulting in lower strain in the tensile reinforcement and a
lower degree of cracking. Alternative variations of metallic anchorage devices were
used by Aridome et al. (1998), The configurations investigated are shown in Figure 2.9.
Although strengthened beams without any anchorage at the underside of the flange were
not tested, the re- searchers reported yielding of the main flexural reinforcement in all
the strengthened beams with steel anchorages. It was also found that the strengthened
beams with angles bolted into the flange reached a higher load than bolting angles into
the web. This has been consistently verified by many researchers.
Chapter 2 – Literature Review of FRP Anchorage Systems in concrete Infrastructure
31
Figure 2.9 – Steel anchorage schemes for strengthening of T-beams in shear; adapted from (Aridome et al. 1998).
2.5.2 Anchorage of FRP through concrete embedment
Embedment of the L-shaped or U-shaped fibers within the flange of the T-beam is a
form of anchorage involving local cutting, breakout, and removal of concrete to the
underside of the beam flanges. The breakouts are typically filled with epoxy resin after
embedment with composite fiber ligatures, as presented in Figure 2.10. Although
lacking the inherent drawbacks of full wrapping because no access is required to the top
of the slab, embedment can be a labour intensive, destructive process, particularly
where a small ligature spacing is required.
Pull-out tests reported by Swiss Federal Laboratories for Materials Science and
Technology (EMPA) (1998) have revealed that a 100 mm embedment is sufficient to
develop 60–80% of the tensile capacity of the FRP, while a 200 mm embedded length is
sufficient to develop the full tensile strength of the FRP. Although these figures show
significant promise, the test ignores the high compressive forces in the direction of the
beam’s length which are present in the flange. These forces may in turn affect the
strength of the anchorage.
Chapter 2 – Literature Review of FRP Anchorage Systems in concrete Infrastructure
32
Figure 2.10 - (a) Typical FRP plate embedded 150mm into beam side with epoxy resin (b) Typical schematic of typical test setup
Lee and Al-Mahaidi (2008) and Lee (2003) conducted large scale experimental
investigations on the shear-strengthening of reinforced concrete T-beams using two L-
shaped shear jackets 40 wide and 1.2 mm thick. The shear jackets were embedded 100
mm into the flange of the beam for suitable anchorage. Photogrammetry was used to
record deformation measurements. Anchor- age failure was initiated at the beam soffit
by an abrupt ripping of a concrete portion at the CFRP bend zone, resulting in
separation failure of the CFRP laps at the beam soffit (Lee 2003). Measurements of
average strains indicated that 5;500–8;884 was achieved prior to the occurrence of
this failure. Because no observable CFRP pull-out from the flange was recorded, it is
difficult to assess the residual capacity of the top embedment anchorage. It is believed
that the use of the rigid L-plates may have been responsible for the initial debonding
due to peeling stresses being introduced at the beam soffit. The use of U-jacketing with
flange embedment would therefore be a more effective method of strengthening.
2.5.3 FRP spike anchors in shear applications
To increase the effectiveness of FRP shear reinforcement applied to T-beams or in
slab/column wall interface configurations, the use of FRP anchors has been proposed
for end anchorage. Typically, a fiber tow made up of braided fibers to form a string is
placed into a predrilled hole in the concrete and filled with adhesive. The fiber ends are
splayed outward in a fan shape and fully bonded to the FRP ligatures with epoxy resin.
A typical representation is shown in figure 2.11.
Chapter 2 – Literature Review of FRP Anchorage Systems in concrete Infrastructure
33
Figure 2.11 - Typical details of FRP spike anchors applied to shear applications
Experimental tests using various configurations have shown that the anchorages are
effective in terms of deformability and strength increase, characteristics which are
dependent on the number of anchorages used (Ceroni et al. 2008). Experimental testing
to determine the improvement from the use of such anchors has been limited to date. In
the context of the anchor pull-out scenario shown in Figure 2.11, experiments have been
conducted to date. Investigations have been carried out by Ozdemir (2005) to determine
the required embedment depth into the concrete to achieve full development of the
anchor under pull-out conditions. Ozdemir determined that there is an effective
embedment depth after which the capacity of the anchor no longer increases. Tests were
conducted using 10–20 MPa concrete with 14–20 mm diameter anchors, and the
embedment depth was suggested as 100 mm. Ozbakkaloglu and Saatcioglu (2009) also
conducted a large number of pull-out tests with 25–100 mm embedment and concluded
that an increase in embedment length results in a decrease in the average bond strength.
This implies that the bond stress distribution decreases with increasing bond length.
Tests and modelling of FRP anchors subjected to pull-out forces have also been
undertaken by Kim and Smith (2009a, b, 2010).
An important characteristic of FRP anchors is the bend that exists between the braided
fiber toe embedded in the concrete and the fanned portion of the anchor in shear
applications. This bend is typically 90 degrees. ACI 440.2R-08 (2008) states that where
fibers wrap around the corners of rectangular cross sections, the corners should be
rounded to a minimum 13 mm radius to prevent stress concentrations in the FRP
system. Specimens tested by Pham and Bayrak (2009) utilized a bend radius ranging
from 0–12 mm and recorded a 23% reduction in anchor strength when no bend radius
was used. Based on previous research by the Japan Society of Civil Engineers (JSCE)
Chapter 2 – Literature Review of FRP Anchorage Systems in concrete Infrastructure
34
(2001), anchors could lose about half of their tensile capacity due to the stress
concentration caused by the anchor bend. Orton et al. (2008) suggested that anchors
with two times the cross-sectional area of the longitudinal CFRP should be used in
practice. Ozbakkaloglu and Saatcioglu (2009) also investigated the effects of inclined
anchors with inclination angles of 0, 15, 30, and 45 degrees. It was found that an
inclination angle of 45 degrees reduced the pull-out load by over 50%. However, no
mention was made of a transitional radius and the system was penalized by high stress
concentrations at the corners, resulting in partial crushing of 20–30 mm deep concrete
under the horizontal compressive stresses transferred by the anchors.
In addition to the joint information provided in the FRP anchor section, the distance of
the anchor from the concrete free edge (closest to the point of load application) was
found to be of importance by Kim and Smith (2009a, b). Kim’s study showed the failure
load to increase the closer the anchor is positioned to the concrete free edge. This
suggests that anchors should be positioned in zones where interfacial shear stresses are
the highest. Also of importance is the stress transfer mechanism from the anchor fan to
the CFRP fiber. According to Kobayashi et al. (2001), if stresses are to be transferred
from one FRP fiber to another using a fan, the fan opening angles should be limited to
less than 90° to limit stress concentrations and prevent premature fracture of the FRP
fiber.
FRP spike anchors have also been successful in strengthening L-shaped concrete
specimens confined with FRP jackets. Karantzikis et al. (2005) concluded that a limited
strength increase is observed in the use of jackets without anchors, regardless of the
FRP thickness used. This was due to poor utilization of the FRP as a result of premature
debonding at the re-entrant corner. Partial depth FRP anchors were found to allow the
jacket to deform substantially and even approach its tensile capacity. Increases in
strength of 20–30% were seen due to the anchors only. The use of full-depth anchors
resulted in increased strength (49% increase due to anchors only) but marginal benefits
in deformability. Further research has demonstrated that FRP jackets and anchors
effectively confine deficient column lap splices and successfully alter the column failure
mode from brittle splice failure to yielding of column reinforcement (Kim et al. 2009).
It was found that increasing the spacing of anchors improved the strength of the splice,
Chapter 2 – Literature Review of FRP Anchorage Systems in concrete Infrastructure
35
while deformation capacity was improved by using a greater number of smaller anchors.
There is currently a lack of available data in which FRP anchors have been applied to
anchor FRP shear fibers, where sufficient measurements were reported. This should be a
focus for future studies.
2.5.4 Evaluation of FRP anchors used to strengthen members in Shear
In order to evaluate the various types of anchorages used to increase the effectiveness of
FRP shear strengthened members, a classification and evaluation approach is adopted
based on the effective strain approach given in (ACI 440.2R-08 2008) section 11.4.1 for
shear strengthened members. The FRP effective strain is used to determine the
anchorage effectiveness factor (kfas), refer equation 2.2.
(2.2) (2.3)
(2.4) (2.5)
Cha
pter
2 –
Lite
ratu
re R
evie
w o
f FRP
Anc
hora
ge S
yste
ms i
n co
ncre
te In
fras
truc
ture
36
Tabl
e co
ntin
ued
on n
ext p
age…
Aut
hor
Des
igna
tion
Com
men
ts
f' c
t ft E f
Sh
ear
k fas
Fa
ilure
Typ
e1
MPa
m
m
GPa
f,m
ax
Con
trol -
no
stre
ngth
enin
g
1.
00
(Ave
rage
) (T
anar
slan
, Mur
at e
t al.
2008
) Sp
ec-2
Si
de b
onde
d CF
RP
31.9
0.
12
231
2000
0.
44
S+C
SF
(Tan
arsl
an, M
urat
et a
l. 20
08)
Spec
-4
U-ja
cket
ing
CFR
P 29
.1
0.12
23
1 16
00
0.37
S+
CSF
(K
halif
a, B
elar
bi e
t al.
2000
) A
-SO
3-2
U-ja
cket
strip
s, 50
@ 1
25m
m
27.5
0.
20
228
4700
1.
41
CSF
(K
halif
a, B
elar
bi e
t al.
2000
) A
-SO
3-4
One
ply
con
tinuo
us U
-jack
et
27.5
0.
20
228
4500
1.
35
CSF
(K
halif
a, B
elar
bi e
t al.
2000
) C-
BT2
O
ne p
ly c
ontin
uous
U-ja
cket
35
.0
0.20
22
8 45
00
1.15
C
SF
(Kha
lifa,
Bel
arbi
et a
l. 20
00)
B-C
W2
Two
plie
s (90
°/0°
) 27
.5
0.30
22
8 27
00
0.99
C
SP
(Kha
lifa,
Bel
arbi
et a
l. 20
00)
A-S
W3-
2 Tw
o pl
ies (
90°/
0°)
19.3
0.
30
228
2300
1.
06
CSP
(K
halif
a, B
elar
bi e
t al.
2000
) A
-SW
4-2
Two
plie
s (90
°/0°
) 19
.3
0.30
22
8 19
00
0.88
C
SP
CFR
P+ M
etal
lic A
ncho
rs
1.76
(A
vera
ge)
(Arid
ome,
Kan
akub
o et
al.
1998
) N
o. 2
4 an
gle
with
thro
ugh
bolt
18.0
0.
12
229
6000
2.
01
FF +
FR
(T
anar
slan
, Mur
at e
t al.
2008
) Sp
ec-3
L-
shap
ed C
FRP
+ St
eel A
ncho
rage
30
.7
0.12
23
1 47
00
1.06
S+
FR
(Tan
arsl
an, M
urat
et a
l. 20
08)
Spec
-5
U-ja
cket
ing
CFR
P +
Stee
l Anc
hora
ge
30.7
0.
12
231
6000
1.
36
S+FR
(T
anar
slan
, Mur
at e
t al.
2008
) Sp
ec-6
L-
shap
ed C
FRP
+ St
eel A
ncho
rage
30
.8
0.12
23
1 47
00
1.06
FF
(T
anar
slan
, Mur
at e
t al.
2008
) Sp
ec-7
Ex
tend
ed U
-Jac
ket C
FRP
+ St
eel A
ncho
rage
30
.6
0.12
23
1 78
00
1.77
FF
(G
alal
and
Mof
idi 2
010)
S-
M-D
U
-jack
etin
g C
FRP
(unb
onde
d) +
Anc
hora
ges
43.0
0.
1723
0 42
00
0.98
S
(Fra
nces
co, R
aghu
et a
l. 20
02)
JS3A
1
ply
CFR
P lig
atur
es +
Anc
hor
20.6
0.16
228
7500
2.
48
FR
(Kha
lifa,
Bel
arbi
et a
l. 20
00)
C-B
T6
Con
tinuo
us U
-jack
et w
ith e
nd a
ncho
r 35
.0
0.20
22
8 63
00
1.61
FF
(F
ranc
esco
, Rag
hu e
t al.
2002
) JS
6A
2 pl
y A
FRP
ligat
ures
+ A
ncho
r 20
.60.
30
117
3400
1.
09
FR
(Fra
nces
co, R
aghu
et a
l. 20
02)
JS5A
2
ply
CFR
P lig
atur
es +
Anc
hor
20.6
0.33
22
8 56
50
2.62
FR
(D
eifa
lla a
nd G
hoba
rah
2010
) TB
1S1
U-ja
cket
ing
CFR
P +
Stee
l Anc
hora
ge
25.6
0.
86
63.6
42
60
1.51
C
SF +
T
(Dei
falla
and
Gho
bara
h 20
10)
TB1S
2 Ex
tend
ed C
FRP
U-J
acke
t + S
teel
Anc
hora
ge
25.6
0.
86
63.6
47
00
1.67
C
SF +
T
(Dei
falla
and
Gho
bara
h 20
10)
TB1S
3 Fu
ll w
rapp
ing
+ St
eel A
ncho
rage
25
.6
0.86
63
.6
7690
2.
73
CSF
(D
eifa
lla a
nd G
hoba
rah
2010
) TB
3S4
Com
bine
d U
-wra
ppin
g an
d Ex
tend
ed
U-J
acke
t +
Stee
l 25
.6
0.86
63
.6
7590
2.
70
CSF
1 A
RS (A
ncho
rage
failu
re a
t sof
fit);
ASF
(Adh
esiv
e se
para
tion
failu
re);
CSF
(con
cret
e se
para
tion
failu
re);
FF (F
lexu
ral f
ailu
re);
FR (f
iber
rupt
ure)
; PFR
(par
tial f
iber
rupt
ure)
; CPO
(Con
cret
e pu
ll-ou
t fai
lure
); PA
SF (p
artia
l adh
esiv
e se
para
tion
failu
re);
S (S
hear
failu
re);
CSP
(Con
cret
e sp
littii
ng);
LR (L
amin
ate
rupt
ure)
; PLR
(Par
tial l
amin
ate
rupt
ure)
; PFR
(Par
tial f
iber
rupt
ure)
; T
( tor
sion
al fa
ilure
of c
oncr
ete)
Cha
pter
2 –
Lite
ratu
re R
evie
w o
f FRP
Anc
hora
ge S
yste
ms i
n co
ncre
te In
fras
truc
ture
37
Tab
le 2
.2 -
CFR
P Sh
ear A
ncho
rage
dev
ices
sum
mar
y
Aut
hor
Des
igna
tion
Com
men
ts
f' c
t ft E f
Sh
ear
k fas
Fa
ilure
Typ
e1
MPa
m
m
GPa
f,m
ax
CFR
P+ E
mbe
dmen
t in
flang
e
4.
27
(Ave
rage
) (L
ee 2
003)
B
eam
0.7
5D
CFR
P L-
strip
s + 1
20m
m E
mbe
dmen
t in
flang
e 31
.1
1.30
13
7.3
8884
4.
81
AR
S (L
ee 2
003)
B
eam
0.6
D
CFR
P L-
strip
s + 1
20m
m E
mbe
dmen
t in
flang
e 30
.9
1.30
13
7.3
7298
3.
97
AR
S (L
ee 2
003)
B
eam
0.5
D
CFR
P L-
strip
s + 1
20m
m E
mbe
dmen
t in
flang
e 31
.6
1.30
13
7.3
7515
4.
03
CPO
C
FRP+
Ful
l wra
p th
roug
h fla
nges
4.
8 (A
vera
ge)
(Hou
lt &
Lee
s 200
9)
B3/3
0/H
/22
Full
Wra
p th
roug
h 45
°hol
es c
ut h
ighe
r int
o fla
nges
22
.3
1.60
12
1.0
6050
4.
09
S (H
oult
& L
ees 2
009)
B4
/30/
G/2
5 Fu
ll W
rap
thro
ugh
45°h
oles
cu
t hi
gher
int
o fla
nges
+ h
oles
24
.6
1.60
12
1.0
7700
4.
88
S (H
oult
& L
ees 2
009)
B5
/30/
C/2
7 Fu
ll W
rap
thro
ugh
45°h
oles
cu
t hi
gher
int
o fla
nges
+ h
oles
26
.7
1.60
12
1.0
9050
5.
43
S 1 A
RS (A
ncho
rage
failu
re a
t sof
fit);
ASF
(Adh
esiv
e se
para
tion
failu
re);
CSF
(con
cret
e se
para
tion
failu
re);
FF (F
lexu
ral f
ailu
re);
FR (f
iber
rupt
ure)
; PFR
(par
tial f
iber
rupt
ure)
; CPO
(Con
cret
e pu
ll-ou
t fai
lure
); PA
SF (p
artia
l adh
esiv
e se
para
tion
failu
re);
S (S
hear
failu
re);
CSP
(Con
cret
e sp
littii
ng);
LR (L
amin
ate
rupt
ure)
; PLR
(Par
tial l
amin
ate
rupt
ure)
; PFR
(Par
tial f
iber
rupt
ure)
; T
( tor
sion
al fa
ilure
of c
oncr
ete)
Chapter 2 – Literature Review of FRP Anchorage Systems in concrete Infrastructure
38
The equation has been based on factors such as: concrete strength (f’c), fiber thickness
(tf), fiber modulus (Ef), depth of FRP ligature (dfv) and maximum fiber elongation
( f,max) . A summary of anchorage data compiled from various researchers is presented
in table 2.2 along with the corresponding anchorage effectiveness factors. The results
are limited by the relatively small number of specimens tested under each experimental
regime and the lack of publications that present strain data of FRP ligatures prior to
failure. In addition, workmanship, material properties, specimen geometry and loading
procedure can affect the accuracy of cross-comparisons between different experimental
programs. The anchorage effectiveness factor ignores these parameters as well as those
not included in the equation.
Complete wrapping though beam flanges has shown the highest anchorage
effectiveness, which is to be expected due to a lack of FRP termination point weakness.
However, this method is labour intensive, involving localised removal of concrete with
the potential for damaging existing reinforcement. It is observed that full wrapping
through beam flanges resulted in the highest observed average anchorage effectiveness
factor of 4.8. Due to the wide scatter of results presented by Hoult & Lees (2009), some
of his data has been omitted from table 2.2 due to flexural failure mode which masked
the performance of the anchorages. The data for flange embedment anchors is currently
limited and more data is required to establish statistical reliability, the anchorage
effectiveness factor of 4.27 is well above the other forms of anchorage. Metallic
anchorages have not shown the same degree of effectiveness in shear applications as for
flexure. When applied to FRP stirrups, they are the least effective form of anchorage
resulting in a standardized anchorage effectiveness factor (kfa)s of 1.64.
Chapter 2 – Literature Review of FRP Anchorage Systems in concrete Infrastructure
39
2.6 Conclusions
The anchorage of externally bonded FRP materials is one means by which higher FRP
strain levels may be achieved prior to failure. The beneficial uses for FRP anchorage
systems are seen to result in achieving higher levels of strengthening using less material
with a more timely installation process. Commonly documented anchor- age solutions
for FRP-to-concrete applications with encouraging results have been presented in this
paper and include (1) FRP U-jackets, (2) FRP spike anchors, (3) mechanically fastened
steel plates and (4) concrete embedment. Published data on the above FRP anchorage
devices was consolidated and presented in terms of an anchorage effectiveness factor in
order to evaluate anchorage efficiency. A framework of the resulting presentation was
given and this will aid future researchers in reporting key measurements.
Metallic anchorages have been demonstrated to be the most effective form of FRP
anchorage devices when applied to flexural FRP. This is the case when using the
maximum fiber elongation prior to failure as the sole evaluation criteria. However,
metallic anchorages require a labour intensive installation process, they are subject to
corrosion, and require regular maintenance. It is recommended that metallic anchorages
be used where a high level of anchorage is required that cannot be achieved by using
non-metallic anchors. When evaluating non-metallic anchors, it was found that inclined
U-jackets were 74% more effective than vertically orientated U-jacket anchors,
resulting in an exceptionally high anchorage efficiency (kfab = 1.36). U-jackets are non-
destructive and easy to install, making them ideal choices for flexurally strengthened T-
beams. In spite of the limited research conducted in the area of FRP anchors, they
have shown good promise. FRP anchors were 46% more effective than vertically
orientated U-jackets and slightly less effective than inclined U-jackets. They can be
applied to both beams and slabs to increase anchorage performance and require no
development along beam sides. FRP anchors require localized drilling, but are relatively
simple to install and are low maintenance. The choice of anchorage specified in practice
will be governed by member geometry and the level of anchorage required.
When considering FRP anchorage devices developed for strengthening in shear and
torsion applications, the use of flange embedment demonstrated the highest overall
Chapter 2 – Literature Review of FRP Anchorage Systems in concrete Infrastructure
40
strain benefit (kfas = 4.27) apart from full wrapping, the limited amount of available test
specimens suggests that more data is required for further verification. Flange
embedment requires localized breakouts for FRP insertion and the destructive nature of
the anchor makes this solution appear unattractive in practical applications. Contrary to
the high level of performance shown for flexural members, metallic anchorages were
found to have the lowest efficiency (kfas = 1.76) when applied to anchor FRP shear
fibers and are therefore not recommended for shear strengthening applications.
Although the improvements in strength due to the anchorage of FRP materials has been
clearly demonstrated, there remains a lack sufficient numerical and experimental data in
the literature to develop extensive databases with statistical reliability that can be used
to develop strength prediction models. It is recommended that future development of
FRP anchorages focus on examining the various anchorage types presented in more
detail. Research should make use of experimental and numeric parametric studies to
inform strength prediction models that may be incorporated into future FRP design
guidelines.
Chapter 3 – Literature Review of FRP-to-Concrete Bond Behaviour
41
3 CHAPTER 3 – LITERATURE REIEW OF FRP-TO-CONCRETE BOND BEHAVIOUR
There have been a number of experimental studies on the bond behaviour between
concrete and FRP (Van Gemert 1980; Taljsten 1994; Chajes et al. 1996; Maeda et al.
1997; Yuan and Wu 1999). Researchers have experimented with alternative test
configurations which have resulted in a variety of failure modes. In this chapter, a
summary of the test setups which have been used to evaluate the strength of FRP-to-
concrete joints is reported. This is followed by an overview of the observed stress
transfer mechanisms between the FRP and the concrete and the influencing parameters.
Lastly, studies on the bond-slip relations and concrete fracture energy are reviewed.
3.1 Test set-ups and failure modes
Several test configurations have been proposed to study the bond between the FRP and
the concrete but no consensus on a standard test procedure has been reached. Chen and
Teng (2001) classified the existing test-setups into the following types: (a) far end
supported double shear tests; (b) near end supported double shear tests; (c) far end
supported double single tests; (d) near end supported single shear tests; (e) beam tests
and (f) modified beam tests.
Of these, far end supported double shear tests and near end supported single shear tests
are the most popular due to their simplicity. In crack-induced de-bonding failures, the
stress state in the critical region of a beam is closely similar to that of a concrete prism
in a near end supported (NES) single shear pull test and the latter serves as a promising
candidate for a standard set-up for determining the FRP-to-concrete bond strength
(Camata et al. 2004; Yao et al. 2005). Figure 3.1 summarises the most common test set-
ups:
Chapter 3 – Literature Review of FRP-to-Concrete Bond Behaviour
42
Figure 3.1 - Different set-ups for shear-lap tests: a) Double pull-pull test; b) Single pull-push test; c) Bending test
It has been observed in both numerical and experimental studies that the test setup used
can significantly affect the experimental results (Yao et al. 2005). In addition, factors
such as: the height of the support block in a NES single shear pull test or weather the
concrete block is reinforced have been found to influence the observed failure mode.
For double or single shear tests, six failure models have been observed by Chen and
Teng (2001). The experiments were based on single and double lap-shear test data
collected for 55 specimens. The following failure modes are listed in the order of their
likelihood of occurrence: (1) Concrete failure, (2) Plate tensile failure including FRP
rupture or steel yielding, (3) Adhesive failure, (4) FRP delamination for FRP-to-
concrete joints (5) Concrete-to-adhesive interfacial failure and (6) Plate-to-adhesive
interfacial failure.
(1) Concrete failure signified that the failure surface was in the concrete a few
millimetres beneath the concrete-adhesive interface (otherwise known as concrete cover
separation failure). A concrete prism may also be pulled out near the loaded end (wedge
failure) and has been reported as the most commonly observed failure mode for near
Chapter 3 – Literature Review of FRP-to-Concrete Bond Behaviour
43
end supported single pull tests. Interfacial failure, between either the adhesive and the
concrete (5) or the adhesive and the plate (6) was less commonly reported. Adhesive
failure (3) was also found to be the least likely due to the availability of high strength
adhesives which preclude other forms of debonding failure.
3.2 Bond transfer mechanism
The following principles apply only to the most commonly reported failure mode of
FRP debonding via (1) concrete cover separation failure (concrete failure). It is
generally believed that failure within the concrete is initiated by the formation of a crack
at or near the plate end, due to the high interfacial shear and normal stresses caused by
the abrupt termination of the plate. Once a crack forms in the concrete (at or near the
plate end), the crack propagates and progresses horizontally, along the level of the FRP,
resulting in the separation of the FRP from the concrete. (5) Concrete-to-adhesive
interfacial failure and (6) Plate-to-adhesive interfacial failure can be considered along
similar lines; however the crack and failure plane shifts respectively to the governing
plane of weakness. Debonding can be monitored by observing the stresses along the
FRP bond length. Shear stress distributions have been found to shift as debonding
progresses and stresses are dispersed further along the laminate length. This trend is
depicted in figure 3.2.
Figure 3.2 – FRP-to-concrete joint typical bond stress
distribution (a) top view (b) strain distribution along FRP and (c) shear stress distribution along FRP (Lee, 2003)
Chapter 3 – Literature Review of FRP-to-Concrete Bond Behaviour
44
As a result of the bond transfer mechanism described above, it has been found that
interfacial shear stress distributions between the FRP and the concrete are distributed
over a certain effective bond length, Le. As a result, current bond strength models
predict a bond length beyond which any further increase in the bond length cannot
increase the anchorage strength. This has been confirmed by many experimental studies
(Chajes et al. 1996; Maeda et al. 1997; Chen and Teng 2001);
3.2.1 Parameters influencing bond strength Previous experimental studies into the behaviour of FRP-to-concrete joints have
isolated several key parameters influencing the bond capacity. Such parameters include:
(1) tensile strength of the concrete surface, (2) compressive strength of concrete, (3)
concrete surface preparation, (4) geometry of concrete section, (5) concrete fracture
energy, (6) FRP stiffness, (7) adhesive stiffness, (8) FRP thickness and width (Van
Gemert 1980; Taljsten 1994; Chajes et al. 1996; Maeda et al. 1997; Yuan and Wu
1999).
Chajes et al. (1996) studied the bond and force transfer mechanism of FRP laminates
bonded to the concrete, using a near end supported single shear test configuration. A
series of tests were performed to investigate parameters such as: surface preparation,
concrete strength and adhesive properties. Three different surfaces were investigated:
(1) no surface preparation, (2) a surface ground with a stone to give smooth finish and
(3) a surface abraded mechanically with a wire wheel to leave the aggregate slightly
exposed. The researchers reported that the concrete surface should be mechanically
abriaded or sandblasted to achieve the best possible bond. Yoshizawa (1996) conducted
a similar study by investigating concrete surfaces prepared by either water jetting or
sandblasting. The investigators reported that the surface treatment by a water jet
produced higher bond strength than the surface treatment by sandblasting. The
conclusions of other research in this area is similar: Surfaces treatments that leave a
rough finish result in more available bond area and increase the shear resistance of the
interface.
Chajes et al. (1996) also investigated the influence of concrete strengths ranging from
25 to 45 MPa on the ultimate bond strength of the joint. The study concluded that if the
Chapter 3 – Literature Review of FRP-to-Concrete Bond Behaviour
45
debonding failure is a few millimetres beneath the concrete-adhesive interface, the
ultimate bond strength will proportional to (f’c)0.5, a relationship which is typically
considered as the shear strength of concrete. However, further studies by Horiguchi
(1997) and Izumo (1998) determined that the bond strength between the FRP sheet and
the concrete surface is a function of the concrete tensile strength and proportional to
(f’c)2/3. The term (f’c)2/3 is related to the tensile strength of concrete reported in JSCE
(1996) as 0.23(f’c)2/3. This relationship was subsequently confirmed and adopted in
models proposed by Khalifa et al. (1998) and ACI 440.2R (2008).
The width ratio of the FRP bonded plate to the concrete member (bf /bc) has been
proven to have a significant effect on the ultimate bond strength of the joint. According
to studies by Chen and Teng (2001), a smaller FRP width with respect to the concrete is
expected to result in a non-uniform stress distribution across the width of the concrete
member, resulting in higher shear stresses in the adhesive at failure attributed to the
contribution of concrete outside the bond area. Taking the above into account, a simple
linear bond strength model was proposed where the width ratio between the FRP plate
and the concrete were related. A smaller FRP width with respect to the concrete resulted
in a higher width ratio coefficient and corresponding ultimate bond strength. This
observation was confirmed by Subramaniam et al. (2007), who investigated 5
alternative FRP widths ranging from 12 to 46mm. The study confirmed the general
experimental trend which revealed that a higher FRP bond strength was achieved with a
decrease in the width of the FRP plate with respect to the concrete.
Nakaba et al. (2001) investigated the effect of FRP stiffness (fiber modulus and
thickness), concrete strength and adhesive thickness. Although the maximum load
reached prior to debond increased with increasing FRP stiffness, the maximum FRP
elongation reached prior to debonding was found to reduce. This was a result of the
higher interfacial shear stresses between the adhesive and the concrete which increase
proportionally with FRP stiffness.
Chapter 3 – Literature Review of FRP-to-Concrete Bond Behaviour
46
3.3 Modelling FRP Debonding
3.3.1 Bond slip models The most common procedures for modelling the FRP-to-concrete interface can be
divided into three main groups summarised by Freddi and Savoia (2008): (1) those
based on stress analysis, where the FRP-to-concrete interface is modelled as a linear
elastic layer and debonding occurs when the interfacial stresses reach the shear strength
of the interface (Malek 1998); Although the method is simple and can be used for bond
slip models and 2D finite element models, it ignores the non-linear load transfer in
mode II softening conditions. (2) Models based on linear elastic fracture mechanics
(LEFM) rely on the assumption of initial cracks or delamination and cannot be applied
in regions where delamination has not occurred (Buyukozturk and Hearing 1998);
Lastly, (3) cohesive crack models have been developed to simulate the fracture process.
Here, debonding is modelled by considering a fictitious crack and through the use of
constitutive relations which consider crack opening and sliding related to cohesive
tractions (Camacho and Ortiz 1996).
Cohesive crack models have become one of the more widely used to study FRP
debonding, where deboning occurs via fracture of the concrete layer beneath the FRP
material. Mode II fracture tends to dominate and numerical models are typically based
on shear stress–tangential slip (bond-slip) interface laws.
Mode II interface laws can be obtained from experimental strain gauge measurements
along FRP reinforcement in a shear test performed up to complete debonding (Freddi
and Savoia 2008). Considering an elastic behaviour for the composite, average shear
stress values between two subsequent strain gauges (or any two points along the
laminate) can be written as a function of the difference in the measured strains, ei, ei+1.
This equates to dividing the force difference between two points by the total area and is
represented in equation 3.1:
(3.1)
Chapter 3 – Literature Review of FRP-to-Concrete Bond Behaviour
47
Where Ef and tf are FRP elastic modulus and thickness; f,i+1 and f,i are FRP strains;
and L is the distance between monitoring points.
In order to define the slip distribution along the FRP plate, the following assumptions
are made:
(1) Perfect bonding (no slip) between plate and concrete at last strain gauge position;
(2) Deformation of concrete specimen far from external cover is negligible with respect
to the FRP counterpart;
(3) Linear variation of strains in FRP plate between two subsequent strain gauges;
(Ferracuti et al. 2007)
The average slip is then calculated as the incremental sum of the FRP extension. This is
expressed in equation 3.2:
(3.2)
In general, the bond–slip curves have a non-linear ascending and descending trends. It
was found that these trends can be approximately described using Popovics’ equation
given by equation 3.3:
(3.3)
Where : and s1 are the maximum bond stress and corresponding slip. The value
(a) controls the slope of the ascending and descending branches of the bond slip curve.
A value of a = 3 was established by (Nakaba et al. 2001). A typical bond slip curve
fitted using experimental data is shown in figure 3.3.
Chapter 3 – Literature Review of FRP-to-Concrete Bond Behaviour
48
Figure 3.3 – Typical bond slip curve Many theoretical models have been developed to predict the bond strengths of FRP-to
concrete bonded joints, generally on the basis of pull test results and eight of them have
been examined in detail by Chen and Teng (2001). For a bond–slip model to provide
predictions to a high level of accuracy, it needs to have an appropriate shape as well as a
correct value for the peak shear stress, corresponding slip and interfacial fracture energy
which is equivalent to the area under the bond–slip curve. The formulations of 5
different bond-slip models are presented in table 3.1 for a typical FRP-to-concrete
bonded joint. From the equations it can be seen that the shapes of the bond-slip curved
vary substantially. Neubauer (1999) proposed a simplistic linear brittle model quite
different from the others, which generally follows non-linear ascending and descending
branches. Such non-linear bond slip behaviour is observed in the majority of
experimental studies and is represented well in Nakaba et al. (2001) .
Cha
pter
3 –
Lite
ratu
re R
evie
w o
f FRP
-to-C
oncr
ete
Bond
Beh
avio
ur
49
Tabl
e 3.
1 –
Sum
mar
y of
pro
pose
d bo
nd sl
ip m
odel
s for
FR
P-to
-con
cret
e jo
ints
Bo
nd S
lip M
odel
N
euba
uer
and
Ros
tasy
, 199
9 N
akab
a et
al,
2001
M
onti
et a
l, 20
03
Savi
oa e
t al,
2003
D
ai a
nd U
eda,
200
3
Asc
endi
ng
Bran
ch, s
so
Des
cend
ing
Bran
ch, s
>
so
0
max
so
0.
065
0.05
1
sf
w
Rem
arks
A
line
ar a
scen
ding
br
anch
with
a
sudd
en d
rop
A si
ngle
cur
ve
A si
ngle
cur
ve
3.5
3.5
Chapter3 – Literature Review of FRP-to-Concrete Bond Behaviour
50
3.3.2 Concrete fracture energy methods To determine the load at which FRP’s debond from the concrete substrate using fracture
mechanics principles, the single most important parameter is the concrete fracture
energy. The fracture energy can be calculated by the area under descending branch of
the bond-slip curve. Fracture mechanics has demonstrated three modes in which a crack
may propagate: Mode I fracture is classified as an opening mode where the tensile
stresses are normal to the plane of the crack. Mode II is a sliding mode where
crack propagation is propelled by shear stresses acting parallel to the plane of the crack
and normal to the crack front. Mode III fracture is classified as a tearing mode
with shear stresses acting parallel to the crack plane and parallel to the crack front.
Research into FRP debonding has proven that despite the fibers being primarily loaded
in shear, the initiation of deboning is still regarded as a mode I fracture in concrete.
Researchers have investigated the use of both modes I and II fracture energy values into
various numerical models and it has been shown that the use of mode I fracture energy
gives strength predictions that result in closer correlations with the test results in the
majority of cases.
Currently available mode I fracture energy correlations are empirical formulations
derived from experimental procedures (Bazant 1998). Consequently, the majority of
fracture energy models are based on aggregate size, concrete strength and water cement
ratio (Van Mier 1997; Trunk and Wittmann 1998; Neubauer and Rostasy 1999; Bazant
and Becq-Giraudon 2002; Ulaga and Vogel 2003; Elsayed et al. 2007; Freddi and
Savoia 2008).
One of the earliest formulations proposed by Bazant (1983) for mode I fracture energy
of concrete was empirically derived, based on notched beam specimens and correlated
with concrete compressive strengths. The formulation is given by:
(7)
Chapter 3 – Literature Review of FRP-to-Concrete Bond Behaviour
51
Where f’ct is the concrete tensile strength, Ec is the elastic modulus of the concrete and
da is the concrete maximum aggregate size.
An alternative model for GIF was proposed in the CEB-FIB model code (CEB-FIB
1990) which was derived based on test data reported in the literature up to the time of
publication:
(3.4)
Van Mier (1997) proposed a simple equation correlating the fracture energy directly to
the concrete compressive strength:
(3.5)
On the other hand, Trunk and Wittmann (1998) investigated the correlation between
fracture energy and the maximum aggregate size which resulted in the corresponding
power function:
(3.6)
A slightly different procedure was derived by Neubauer (1999) based on the results of
single and double shear tests. The proposed formulation took into account the
geometrical relation between the width of the FRP plate, bf and the width of the
concrete member, bc and the concrete tensile strength:
(3.7) (3.8)
Where cf is an empirical constant containing all secondary effect reported as 0.202.
Bazant and Becq-Giraudon (2002) showed that the concrete fracture energy can be
approximately predicted from the standard compression strength, maximum aggregate
size, water–cement ratio, and aggregate type (river or crushed). The model was
Chapter 3 – Literature Review of FRP-to-Concrete Bond Behaviour
52
developed based on statistical calibration with an extensive database built from the
literature, where o equals 1.12 for crushed aggregates:
(3.9)
The influence of adhesive shear stiffness (Ga) and thickness (ta) on the overall
interfacial fracture energy was investigated by Dai (2003). Based on their test results the
following expression was derived:
(3.10) (3.11)
It is apparent that the above formulations based on seven different procedures used to
determine concrete fracture energy will produce a range of values. The above summary
is presented to provide a brief review of a range of available formulations for fracture
energy predictions. These will be compared later in chapter 7 using the material
properties adopted in the experimental program and used as the basis for sensitivity
studies within the construction of calibrated numerical models.
3.3.3 Bond strength models
Several anchorage strength models proposed in recent years have been summarised by
(Chen and Teng 2001; Sayed-Ahmed 2009) where models were divided into three
categories: (1) empirical models derived from experimental data; (2) models based on
fracture mechanics and (3) design proposals that make use of some simple assumptions.
One of the earliest experimentally based models which also used simplistic assumptions
was introduced by Van Gemert (1980). Based on an examination of the shear stresses in
a double shear test, the tensile force was found to decay towards the anchored end of the
plate. The model also assumed a triangular stress distribution with linear ascending and
descending branches for the bond-slip curve and considered the effects of FRP laminate
width, effective length and concrete characteristic tensile strength. However,
experimental evidence has shown that this assumption is not correct for the FRP-to-
Chapter 3 – Literature Review of FRP-to-Concrete Bond Behaviour
53
concrete bond surface as the bond-slip behaviour tends to follow non-linear ascending
and descending branches.
A fracture mechanics based model was introduced by Holzenka¨mpfer (1994) which
investigated the bond strength between a steel plate and concrete. The model was
derived on the basis that the concrete fracture energy can be derived as a function of the
concrete tensile strength. Neubauer (1997) subsequently modified this model so that it
applies to both FRP and steel plates and the resulting formulation is presented in table
3.2. Using non-liner finite element analysis, Taljsten (1994) developed a similar model
which considered the influence of concrete fracture energy and FRP stiffness with
respect to the concrete. The same model was later expanded upon by Yuan and Wu
(1999) to include the width ratio of the FRP bonded plate to the concrete member.
Existing bond strength models have been reviewed and assessed by Chen and Teng
(2001) through a comparison with experimental data gathered from literature. After
assessing the strengths and weaknesses of each modal, a new simplified fracture
mechanics based model was introduced capturing all of the main features of anchorage
behaviour, including concrete cylinder strength, effective bond length, FRP stiffness
and with width ratio between the FRP and the concrete.
An empirically derived formulation that considers the effect of effective bond length
and relates the average bond strength to the FRP stiffness and strain gradient was
proposed by Maeda et al. (1997). The effective bond length was derived based on
observations that the no proportional strength increase was observed for bond lengths
over 100mm. Khalifa et al. (1998) proposed a variation of this model by including the
effect of concrete strength by multiplying the maximum shear stress by the term (f’c)2/3.
In the design model proposed by Chaallal (1998), it was assumed that the bond behaves
as a Mohr-Coulomb material which was based on earlier studies by Varastehpour
(1996). Chen and Teng (2001) summarised that for shear strengthening, the model
assumed that the maximum shear stress is twice the average stress and does not exceed
the Mohr-Coulomb strength equation given by Varastehpour (1996).
Chapter 3 – Literature Review of FRP-to-Concrete Bond Behaviour
54
Model Proposed Formulations
(Van Gemert 1980).
(Neubauer 1997)
(Taljsten 1994)
(Yuan and Wu 1999)
(Maeda et al. 1997)
(Khalifa et al. 1998)
(Chaallal 1998)
by (Chen and Teng 2001)
(JCI 2003)
(Yang 2001)
Table 3.2 – Summary of proposed bond strength models for FRP-to-concrete joints
Chapter 4 – Experimental Investigation into FRP Anchorage Systems utilising a
Mechanically Strengthened Substrate
55
4 CHAPTER 4 - EXPERIMENTAL INVESTIGATION INTO FRP ANCHORAGE SYSTEMS UTILISING A MECHANICALLY STRENGHTNED SUBSTRATE
4.1 Introduction
It has been demonstrated that failure of concrete structures retrofitted with FRP usually
occurs by debonding of the FRP from the concrete substrate. To prevent this type of
failure, national standards and design guidelines impose strict limitations on the
allowable strain level in the composite material which may be safely utilised in design.
In order to achieve acceptable levels of FRP-to-concrete contact bond stresses,
offsetting the likelihood of debond, allowable strains are further limited in cases where a
higher degree of strengthening is required. In such cases, design guidelines can limit the
FRP material strain to levels as low as 10-25% of the FRP ultimate tensile strength
(UTS). In practice these limitations result in severe underutilisation of the FRP material
properties reducing economy.
Current literature and design guidelines recognise the benefits of FRP anchorage
systems to increase the bond strength of the FRP to concrete connection. Research has
shown that anchoring the ends of the FRP plates or sheets results in a significantly
higher bond stress being reached before delamination occurs. When sufficiently
anchored, the FRP material strain at failure can approach its ultimate strain at rupture.
Design guidelines such as (ACI 440.2R-02 2002) recognise the benefits of anchorage
systems and permit designers to utilise a higher FRP strain provided that the anchorage
device is backed up by sufficient experimental testing (Refer section 11.4.1.2).
There have only been limited attempts to investigate FRP anchorage measures and
many remain to be quantified. This chapter investigates a new type of FRP anchorage
solution which relies on increasing the mechanical properties of the concrete substrate
to which the FRP is bonded to over a nominal anchorage length.
Chapter 4 – Experimental Investigation into FRP Anchorage Systems utilising a
Mechanically Strengthened Substrate
56
4.2 Specimen Design
4.2.1 The Mechanically Strengthened Substrate Anchor
The efficiency of any anchorage system can be enhanced through an improvement of
the mechanical properties of the concrete substrate over the length of the anchorage
resulting in a higher FRP-to-concrete bond strength. In practice, where the pull off
strength of the concrete substrate is less that the minimum requirements, it is sometimes
possible to improve the pull off bond by impregnating the concrete with a very low
viscosity resin. However, improvement by this method will only occur when the
substrate is porous (BBR ISO 9001 2002). This method also remains to be quantified.
The proposed solution relies on the introduction of a mechanical chase cut into the
concrete over the length of the end anchorage zone. The chase is to be filled with epoxy
resin, prior to bonding the FRP laminate over the prepared surface. The purpose of the
chase is to prevent the critical mode of debond which naturally occurs a few millimetres
beneath the concrete/adhesive interface. It utilises the superior mechanical properties of
the epoxy to distribute the stresses over a larger area and depth within the concrete
prism.
The proposed concept has wide application but was developed specifically for
combined shear/torsional strengthening of box girder bridge webs to be utilised at the
web/flange connection. In addition to the proposed concrete chase a N24 reinforcement
bar was installed within the chase to be embedded into the underside of the bridge deck
to anchor the forces from the ends of the vertical laminates on the outside webs into the
underside of the bridge deck. The purpose of the bar was to augment the amount of steel
reinforcement in the web-flange joint. The latter was found to be inadequate for the
increased loading on the bridge. Due to the method of FRP delamination observed later
in the testing is it stipulated that this reinforced bar does not actively contribute to the
enhancement of the FRP laminate anchorage. It is expected that the omission of this bar
in future testing will not degrade the improvement in FRP-to-concrete bond strength
observed.
Chapter 4 – Experimental Investigation into FRP Anchorage Systems utilising a
Mechanically Strengthened Substrate
57
A full scale test set-up was designed using materials properties prevalent on site for a
comparative study. The factors which were considered in the current test program
included the thickness of the concrete prism tc, FRP bond length Lf and the width ratio
between the FRP strip and the concrete prism bp/bc. Bond strength models developed by
(Ta¨ljsten 1994; Yuan and Wu 1999) have shown that the thickness of the concrete
prism can significantly affect the stress distribution within the specimen. Using the
model proposed by (Yuan and Wu 1999), it was found that a concrete thickness, tc, of
250mm exhibited a 1.0% reduction in bond strength when compared to a 500mm
thickness. This illustrated that an adopted prism depth of 250mm was sufficient to not
adversely influence the overall bond strength of the specimen. It was expected that the
member thickness prevalent in site applications was to be greater than 250mm.
Current bond slip models proposed by researchers stipulate that the width ratio’s of the
FRP to concrete prism (bp/bc) can have some influence on the ultimate strength of a
specimen (Yuan and Wu 1999; Yuan et al. 2001; Pham and Al-Mahaidi 2006). Based
on proposed models by Chen and Teng (2001), it is anticipated that a 10% reduction in
bond strength would be incurred as a result of adopting a prism width of 300mm as
opposed to 500mm.
4.3 Test Preparation and Material properties
Reinforced concrete blocks of dimension 250mm x 300mm x 600mm were used for the
construction of 3 test specimens. A control specimen formed the basis for comparison
together with 2 anchorage specimens which formed the first stages of the experimental
program. Blocks were reinforced nominally with N16-200 each face to replicate the
existing reinforcement present in the webs of the bridge. Table 4.1 summarises the test
specimens constructed for the control specimen and anchorage type 1.
Anchorage Type Ref Anchorage
Description 0 WG9 Control 1 1
WG1 Mechanical chase WG2 Mechanical chase
Table 4.1 - Summary of test specimens constructed in experimental program
Chapter 4 – Experimental Investigation into FRP Anchorage Systems utilising a
Mechanically Strengthened Substrate
58
4.3.1 Control Specimen
The control test consisted of 1 No. 120mmx2mmx1000mm laminate strip bonded to the
surface of the concrete block with a bond length of 500mm (refer figure 4.1). The face
of the concrete block was cleaned, sandblasted to achieve a profile similar to 60 grit
sand paper prior to surface preparation and application of the laminate strip. This was
followed by curing which occurred in a temperature controlled chamber of 50°C for a
period of 48 hrs, then further curing at room temperature (22°C) for a further 72 hours
prior to testing. The accelerated curing was necessary due to tight deadlines in the
experimental program and pressures from industry for experimental results. Adhesion
tests were carried out on additional adhesion samples to verify laminate adhesion prior
to testing.
Figure 4.1 - Control specimen geometry (WG9) configuration of strain gauges;
4.3.2 Anchor Type 1
Torsion in a box girder typically results in a shear flow through the outer webs of the
section. The shear resistance of the outer webs and the girder soffit are both key factors
to ensure that the section has adequate strength to resist the applied torsion. The high in
plane rigidity of bridge decks mean that they seldom need strengthening in practice,
however it is of paramount importance that the web-flange connection contain adequate
reinforcement to adequately develop the tensile stresses induced by torsion into the
bridge deck. For this purpose, a N24 reinforcement bar was installed within the
Chapter 4 – Experimental Investigation into FRP Anchorage Systems utilising a
Mechanically Strengthened Substrate
59
mechanical chase to be embedded into the underside of the bridge deck to anchor the
forces from the ends of the vertical laminates on the outside webs into the underside of
the bridge deck. A by-product of the mechanical chase was an enhancement of the
substrate properties over the length of the anchorage resulting in a higher FRP-to-
concrete bond strength. The construction process involved a 40mm x 40mm x 500mm
chase was cut into the 300mm wide side of the concrete block for the anchorage
specimens. A primer coating was applied to entire surface of concrete block (including
chase) prior to any bonding. An N24 deformed reinforcement bar was then bonded into
the chase using laminate adhesive. In addition, glass fiber fabric (120mm x 400mm)
was applied centrally over reinforcement bar (to prevent galvanic corrosion of the
reinforcement) which is depicted in figure 4.2.
Laminate adhesive was then applied to the underside of the laminate strip and centrally
on the prepared concrete surface. Adhesive was applied to both surfaces using a profiled
template to ensure accurate application (peak thickness of 1.5mm in centre of adhesive
strip). A 120mm x 2mm x 1000mm (210GPa) laminate strip was applied centrally to the
concrete block. Tables 4.2 and 4.3 summarise the material properties of the FRP and
adhesive used to construct the specimens. With the exception of the FRP laminate
whose properties were determined by experimental testing, the properties of the
adhesive were based on the manufacturers specifications. It should be noted that the
curing temperatures provided in table 4.3 were the curing times at different
temperatures provided by the manufacturer and not the temperatures at which the
specimens were cured at.
Properties Bidirectional
FRP (±45°)
FRP
Laminate
Unidirectional
FRP
Bidirectional
GFRP (±90°) Units
Tensile Strength 3.79 3.3 3.8 3.4 GPa Tensile Modulus 230 185 240 73 GPa Ult. Elongation 2.1 1.4 1.55 4.5 %
Density 1.8 1.56 1.7 2.6 g/cm³Thickness 0.55 2 0.235 0.067 mm
Width 120 300 mm
Table 4.2 - FRP Properties data
Chapter 4 – Experimental Investigation into FRP Anchorage Systems utilising a
Mechanically Strengthened Substrate
60
Properties Laminate
Adhesive Saturant Primer Units
Resin Type Epoxy Epoxy Epoxy - Specific Gravity 1.8 1.12 1.08 - Glass Transition >65 - - °C
Modulus of 10 >3.0 0.7 GPa Lap Shear Strength >17 - - MPa Bond (to Concrete) >3.5 >3.5 >3.5 MPa
Tensile Strength 32 >50 >12 MPa Compressive >60 >80 - MPa
Flexural Strength >35 >120 >24 MPa Full cure
at:
25°C 7 7 0.208 Days 40°C 3 - 0.125 Days
Table 4.3 - Adhesives and Saturant Properties data
The laminate strip was pressed down onto the concrete block using a special profiled
tool to ensure accurate thickness of adhesive between concrete surface and laminate
strip and central placement of laminate. Excess adhesive was cleaned up from the
concrete and laminate surface. Specimens cured in a manner similar to the control
specimen. The construction process for anchor type 1 is summarised in figure 4.3. All
samples were tested in a Baldwin universal testing machine.
Chapter 4 – Experimental Investigation into FRP Anchorage Systems utilising a
Mechanically Strengthened Substrate
61
Figure 4.2 - Anchorage type 1 specimen geometry (WG1 & WG2) (a) configuration of strain gauges; (b) chase details and installation of N24 reinforcement bar (c) section through chase.
(a) (b) (c)
Figure 4.3 - Construction process of Type 1 Anchorage Specimen; (a) surface of concrete block coated with MBRACE primer and centralisers for N20 reinforcement bar located within chase; (b) profiling of laminate adhesive (as per manufacturers specification) to the surface concrete block over reinforcement bar; (c) specimen curing at an elevated temperature of 41°C.
Chapter 4 – Experimental Investigation into FRP Anchorage Systems utilising a
Mechanically Strengthened Substrate
62
4.3.3 Experimental Setup
Many alternative experimental set-ups have been used by researchers for determining
the FRP-to-concrete bond strength. Of these far end supported double shear tests and
near end supported single shear tests are most popular due to their simplicity (Camata et
al. 2004; Yao, Teng et al. 2005). In crack-induced de-bonding failures, the stress state in
the critical region of a beam is closely similar to that of a concrete prism in a near end
supported (NES) single shear pull test and the latter serves as a promising candidate for
a standard set-up for determining the FRP-to-concrete bond strength (Camata et al.
2004; Yao, Teng et al. 2005). On this basis the experimental design used in this study
was based on the NES single pull test configuration.
A test rig was constructed to ensure each specimen was able to be securely fixed to a
Baldwin Universal testing machine, the schematics of which are presented in figures 4.4
and 4.5. The test rig was bolted down to the moving lower platform of the testing
machine which clamped the specimen into place. The test rig was constructed using
30mm thick steel plates. A back plate, 600mm high was welded at the rear of the test rig
with 9 No M12 bolts placed across the face of the rear plate to prevent any movement of
the concrete specimens during loading. After the initial series of tests an additional steel
plate was place at the font the specimens (at the base) and bolted to the rear vertical
plate. This prevented any forward movement of the concrete specimens that may occur
during loading.
Once the specimens were centrally located within the testing rig, a rigorous cross
checking program was implemented to ensure the verticality of the test laminate strip. A
spirit level was used to check the verticality of the laminate strip, with shims being used
to create a vertical test specimen. This procedure was cross checked by two independent
people.
Chapter 4 – Experimental Investigation into FRP Anchorage Systems utilising a
Mechanically Strengthened Substrate
63
(a) (b)
Figure 4.4 - Specimen testing rig details (a) configuration of test rig (front view); (b) configuration of test rig (side view)
(b) (b)
Figure 4.5 - Specimen testing rig clamped to Baldwin testing machine (a) configuration of test rig (front view); (b) configuration of test rig (rear view)
Chapter 4 – Experimental Investigation into FRP Anchorage Systems utilising a
Mechanically Strengthened Substrate
64
4.3.4 Instrumentation and loading procedure
Strain and load results were obtained from surface mounted strain gauges and a 3D non-
contact measuring technique based on image correlation photogrammetry (GOM mbH
2005).
A series of 7 strain gauges, from G1 to G7, were attached to the surfaces of FRP plates.
G1 and G2 were installed to monitor any bending in the FRP plate during testing
indicating the presence of tilting. G1 was placed at the back of the laminate and G2 at
the front at the same location. The specimens were tested under displacement control of
0.0167mm/s until beyond de-bonding of the FRP from the concrete specimen.
The 3D photogrammetry measurements were taken using a pair of high resolution,
digital CCD cameras. The image correlation system called ARAMIS by gom optical
measuring techniques (GOM mbH, 2005) was used to acquire the data. A measuring
step of 10 seconds was used between recording intervals. 3D image correlation
software analyses the deformation of a random or regular pattern pixels with good
contrast which is applied to the surface of the specimen and recorded by the CCD
cameras for processing.
4.4 Experimental Results
4.4.1 Quality control tests
Quality control tests consisted in the testing of concrete, adhesive and FRP properties to
determine the actual material properties used in the experiments.
4.4.1.1 Compression strength testing
A total of 6 concrete cylinders were tested to assess the concrete compressive strength
were performed in accordance with AS 1012.9 (1999). After 53 days curing at room
temperature, the average compressive strength of the concrete was 62MPa.
Chapter 4 – Experimental Investigation into FRP Anchorage Systems utilising a
Mechanically Strengthened Substrate
65
4.4.1.2 Pull-off adhesion testing
In order to verify the correctness of surface preparation, concrete tensile strength and
mixing of adhesives, pull off testing was carried out according to I.S. EN 1542 (1999).
Three adhesion tests were performed on the TYFO BCC ±45° fabric with MBRACE
solvant and one additional test was performed on the MBRACE laminate with
MBRACE laminate adhesive. The following procedure was used to conduct the pull-off
testing:
Surface preparation: Surface preparation was performed in the same manner used prior
to application of FRP – which consisted of sandblasting, water jetting and application of
a primer.
Core drilling: A diamond core bit was used to drill 50mm (internal diameter) cylinders
through all FRP and adhesive materials, 5mm deep into the concrete, with an axis of 90
degrees to the surface. The drilling was carried out in order to isolate the area under the
dolly from the surrounding concrete – in order to induce failure within the 50mm
cylinder.
Applying the dolly: After appropriate cleaning of the aluminium dollies using: abrasive
paper and degreaser, the adhesives were prepared according to manufacturer’s
specifications an even layer applied to the dolly and bonded to the centre of the 50 mm
cores.
Applying the load: A DeFelsko adhesion tester, which is shown in figure 4.6 was used
to apply load continuously at an even rate of 0.05 MPa/s until failure occurred.
Test results: The results showed that in all cases failure occurred within the concrete.
Table 4.4 summarises the results obtained for the pull-off tests.
Chapter 4 – Experimental Investigation into FRP Anchorage Systems utilising a
Mechanically Strengthened Substrate
66
Sample No. Material Pull-off strength (MPa)
1 TYFO BCC ± 45° fabric 3.5 2 TYFO BCC ± 45° fabric > 4.0 3 TYFO BCC ± 45° fabric > 4.0 4 MBRACE Laminate 3.6
Table 4.4 - Adhesion test results on TYFO BCC bidirectional fabric and MBRACE laminate strip.
Figure 4.6 - Adhesion testing and pressure gauge reading from test (TYFO BCC ±45° fabric) showing failure within concrete.
4.4.1.3 FRP Laminate properties
The tensile strength and elastic modulus of the FRP laminates were verified using three
laminate coupon tests. FRP composite elastic modulus was determined using testing
procedures in accordance with (ASTM: D 3039 2000). Based on the testing of three
samples a mean elastic modulus of 185GPa was recorded, compared to the
manufacturer’s value of 210GPa.
4.4.2 Failure modes
The control specimen failed by separation of the composite plate from the concrete
block at the interface between the concrete and the adhesive, as shown in figure 4.7. The
failure mode highlights that the interface between the adhesive and the concrete was the
Chapter 4 – Experimental Investigation into FRP Anchorage Systems utilising a
Mechanically Strengthened Substrate
67
weakest plane in the bond line, probably relevant to the high strength of the concrete
substrates. This mode of failure was mitigated in anchorage type 1, as the failure plane
shifted from the concrete-adhesive interface (as observed in the control specimen) to the
adhesive-FRP interface. As a result, the majority of the FRP plate was left exposed,
with no concrete or epoxy bonded to it, as depicted in figure 4.8. The introduction of the
mechanical chase clearly increased the bond strength between the adhesive and the
concrete by an increase of available bond area between the adhesive and the concrete
and the subsequent transfer of stresses within the adhesive to a deeper level within the
concrete.
(a) (b) (c)
Figure 4.7 - Failed Control Sample (WGB9) (a) complete debonding of laminate from concrete surface; (b) concrete surface post debonding of laminate (c) de bonded laminate strip; (d) real time load, strain and ARAMIS photogrammetry recordings during testing phase.
Chapter 4 – Experimental Investigation into FRP Anchorage Systems utilising a
Mechanically Strengthened Substrate
68
(a) (b) (c)
Figure 4.8 - Testing of WGB1 (a) specimen ready for testing; (b) concrete rupture at adhesive concrete interface; (c) debonded laminate strip.
4.4.3 Tilt
In practical pull tests, there may be a small unintended offset in the position of the
load (Yao et al. 2005). The result of any eccentricity in load application can result in a
localised bending effect at the top of the specimen and the likely hood of premature
delamination. Detection and monitoring of any eccentricity has been considered in test
measuring and instrumentation through the installation of strain gauges G1 and G2 at
the front and back of the laminate. The degree of tilting can be determined from the
variation in strains between these two gauges. As shown in figure 4.10, the control
specimen has shown some deviation in strain between gauges G1 and G2 indicating the
presence of tilting. Since G2 shows a higher strain than G1 the bending is expected to
produce push/pull (compressive/tensile peeling stresses) along the length of the concrete
block.
The detrimental effect of eccentricity within subsequent specimens was mitigated by the
use of clamping devices within the test setup, as depicted by the front plate shown in
figure 4.4. Each specimen was tensioned to 25kN to verify accuracy of specimen
mounting. Gauges G1 and G2 were compared to ensure variation between the gauges
were within an acceptable tolerance (±10% of each other), thus ensuring uniform
Chapter 4 – Experimental Investigation into FRP Anchorage Systems utilising a
Mechanically Strengthened Substrate
69
tensioning of laminate during testing. If readings were not within tolerance, the
specimen was unloaded and re-aligned.
4.4.4 FRP strain distributions
In tables and figures which follow reference is made to AR (Photogrammetry) and SG
(strain gauge). These refer to the two data acquisition techniques used in the
experimental programme.
FRP elongation along the length of the laminate are reported in figure 4.9 for both
control and anchored specimens when subjected to different levels of loading.
(a) (b)
(c)
Figure 4.9 - Strain vs distance along Laminate; (a) Control specimen (WG9); (b) Type 1 - Anchorage specimen (WG1) ; (c) Type 2 - Anchorage specimen (WG2)
0
500
1000
1500
2000
2500
3000
0 100 200 300 400 500 600
Micro
strain
()
Distance From Gauge G1 (mm)
60kN(AR)80kN(AR)90kN(AR)96.6kN(AR)60kN (SG)90kN (SG)80kN (SG)96.6kN (SG)
0500
100015002000250030003500400045005000
0 50 100 150 200 250 300 350 400
Micro
strain
()
Distance From Gauge G1 (mm
50kN(AR)100kN(AR)150kN(AR)194kN(AR)50kN (SG)100kN (SG)150kN (SG)194kN (SG)
0500
100015002000250030003500400045005000
0 50 100 150 200 250 300 350 400
Micro
strain
()
Distance From Gauge G1 (mm
STAGE_1.AR/SG.WG2
50kN(AR)
100kN(AR)
150kN(AR)
198.2kN(AR)
50kN (SG)
100kN (SG)
150kN (SG)
198.2kN (SG)
Chapter 4 – Experimental Investigation into FRP Anchorage Systems utilising a
Mechanically Strengthened Substrate
70
Examining figures 4.9 (a)-(c), the photogrammetry data indicates a drop in strain level
at approximately 50mm from strain gauge G1, which corresponds to the edge of the
concrete block. This behaviour can only be observed from the photogrammetry results
as no strain gauge was positioned at this location for correlation. The control specimen
shows a lower drop in strain at this location which could be due to the observed tilt
causing localised bending at the edge of the loaded face where the FRP transitions
between un-bonded and bonded sections. The slope of the strain vs distance curves
indicate a gradual flattening and loss of gradient with increased loading, indicating the
progression of debonding as strain is dispersed further along the FRP bond line.
Figure 4.10 - Load vs strain distribution, control specimen (WG9);
(a)
0
10
20
30
40
50
60
70
80
90
100
0 500 1000 1500 2000 2500 3000
Load
(kN)
Micro strain ( )
G1 (SG)
G2 (SG)
G3 (SG)
G4 (SG)
G5 (SG)
G6 (SG)
G7 (SG)
G1 (AR)
G2 (AR)
G3 (AR)
G4 (AR)
G5 (AR)
G6 (AR)
G7 (AR)
0
20
40
60
80
100
120
140
160
180
200
0 500 1000 1500 2000 2500 3000 3500 4000 4500 5000
Load
(kN)
Micro strain ( )
G1 (SG)
G2 (SG)
G3 (SG)
G4 (SG)
G5 (SG)
G6 (SG)
G7 (SG)
G1 (AR)
G2 (AR)
G3 (AR)
G4 (AR)
G5 (AR)
G6 (AR)
G7 (AR)
Chapter 4 – Experimental Investigation into FRP Anchorage Systems utilising a
Mechanically Strengthened Substrate
71
(b) Figure 4.11 - Load vs strain distribution (a) Type 1 - Anchorage specimen (WG1); (b) Type 1 - Anchorage specimen (WG2)
Figures 4.10 and 4.11, which summarise the load vs strain data at each strain gauge
location, all show good correlation between photogrammetry and strain gauge
measurements.
Progressive debonding is evident in both specimens by the gradual reduction in gradient
of the curves with increasing loading. Sudden debond can be observed by the large
increase in strain at a sustained level of load, which is most apparent in figure 4.10
(control) at 85 kN. The anchorage specimens (type 1) did not exhibit the sudden form
of debonding observed in the control specimen and instead, favoured a more gradual
debonding failure at a much higher load. A comparison of the curves for the control
specimen and anchorage specimen yields the following observations:
Load - strain relations for the anchored specimen (type 1) show a steeper slope which
is indicative of a stiffer substrate compared to the control sample. The failure plane
between the concrete and adhesive was propelled by the presence of mirco cracking
within the concrete during testing – which is believed to have caused the marked
reduction in stiffness of the unstrengthened concrete substrate. This behaviour was
less apparent in anchorage type 1, as the strengthened substrate caused the failure
plane to shift and initiate between the laminate and the adhesive.
0
20
40
60
80
100
120
140
160
180
200
0 500 1000 1500 2000 2500 3000 3500 4000 4500 5000
Load
(kN)
Micro strain ( )
G1 (SG)
G2 (SG)
G3 (SG)
G4 (SG)
G5 (SG)
G6 (SG)
G7 (SG)
G1 (AR)
G2 (AR)
G3 (AR)
G4 (AR)
G5 (AR)
G6 (AR)
G7 (AR)
Chapter 4 – Experimental Investigation into FRP Anchorage Systems utilising a
Mechanically Strengthened Substrate
72
Anchor type 1 strain readings taken at a distance of 150-250mm away from the
loaded face (corresponding to gauges G5-G7) indicated a more efficient load
transfer, resulting in a higher strain achieved prior to the occurrence of partial
debonding. However, the control specimen showed evidence of a sudden increase in
strain at the same location (for gauges G5-G7), which is indicative of sudden debond.
As a result, the strengthened substrate increased the ductility of the joint in addition
to its strength.
The anchorage specimen showed a significant improvement in both the maximum
load and strain reached prior to failure, thus achieving a load level of 194.4-198.5
kN, which was almost twice the load reached by the control specimen. The
enhancement in strength provided as a result of the mechanical chase increasing the
strength of the substrate, was a 95-100% increase in ultimate load capacity and a 83-
93% increase in the maximum strain achieved prior to failure. The maximum failure
loads and strains reached prior to debonding are summarised in table 4.5.
Reference Failure Load (kN)
Max Elongation ( )
GA AR Control Specimen WG9 99.6 2535 2706
Anchorage Type 1 WG1 194.4 4640 4434 WG2 198.5 4881 4733
Table 4.5 - Load/Elongation results summary (WG1, WG2 & WG9)
4.4.5 Experimental bond slip curves
Strain measures along FRP laminate at different loading levels were used to calculate
shear stress–slip data. Considering an elastic behaviour for the composite, the average
shear stress values between two subsequent strain gauges (or any two points along the
laminate) can be written as a function of the difference of measured strains, ei, ei+1. This
equates to dividing the force difference between two points by the total area and is
represented in equation 4.1:
(4.1)
Chapter 4 – Experimental Investigation into FRP Anchorage Systems utilising a
Mechanically Strengthened Substrate
73
Where Ef and tf are FRP elastic modulus and thickness; f,i+1 and f,i are FRP strains;
and L is the distance between strain gauges.
In order to define the slip distribution along the FRP plate, the following assumptions
are made:
(1) Perfect bonding (no slip) between plate and concrete at last strain gauge position;
(2) Deformation of concrete specimen far from external cover is negligible with respect
to FRP counterpart;
(3) Linear variation of strains in FRP plate between two subsequent strain gauges;
(Ferracuti, Savoia et al. 2007)
The average slip is then calculated as the incremental sum of the FRP extension. This is
expressed in equation 4.2:
(4.2)
In general, the bond–slip curves have a non-linear ascending and descending trends. It
was found that these trends can be approximately described using Popovics’ equation
given by:
(4.3)
Where : and s1 are the maximum bond stress and corresponding slip. The value
(a) controls the slope of the ascending and descending branches of the bond slip curve.
A value of a = 3 was established by (Nakaba, Kanakubo et al. 2001).
Table 4.6 demonstrates the effect of the 40x40mm concrete chase on the maximum
bond strength achieved. A strength gain of 118% in bond stress is seen as a result of
introducing the chase anchor.
Chapter 4 – Experimental Investigation into FRP Anchorage Systems utilising a
Mechanically Strengthened Substrate
74
Reference Max Bond Stress (mPa)
Corresponding Slip (so)
GA AR GA AR Control Specimen WG9 5.2 5.1 0.058 0.62
Anchorage Type 1 WG1 11.3 11 0.2 0.2 WG2 17.5 10.9 0.2 0.23
Table 4.6 - Max Bond stress and corresponding slip results summary (WG1, WG2 & WG9) at location 125mm away from concrete free edge.
(a) (b)
(c)
Figure 4.12 - Bond-slip curves (a) Control specimen (WG9) with fitted curve following Popovics equation; (b) Type 1 - Anchorage specimen (WG1)
Figure 4.12 presents the bond-slip curves obtained from both strain gauge and
photogrammetry measurement techniques. A Popovics trend line has been fitted to each
curve using the measured peak bond stress and corresponding slip obtained from each
data acquisition technique. The peak bond stresses determined using the two methods
show an acceptable level of correlation. Photogrammetry measurements estimate +10%
0
1
2
3
4
5
6
0 0.05 0.1 0.15 0.2 0.25 0.3 0.35
Bond
Stress
(MPa
)
Slip (mm)
Popovics
175mm (ARAMIS)
175mm (GAUGE)125 mm
125 mm
0
2
4
6
8
10
12
0 0.05 0.1 0.15 0.2 0.25 0.3 0.35
Bond
Stress
(MPa
)
Slip (mm)
175mm (ARAMIS)
175mm (GAUGE)
125 mm
125 mm
02468
101214161820
0 0.05 0.1 0.15 0.2 0.25 0.3 0.35
Bond
Stress
(MPa
)
Slip (mm)
STAGE_1.AR/SG.WG2
175mm (ARAMIS)
175mm (GAUGE)
125 mm
125 mm
Chapter 4 – Experimental Investigation into FRP Anchorage Systems utilising a
Mechanically Strengthened Substrate
75
higher bond stresses and +17.2% higher slips recorded. The softening branches of the
two bond slip curves follow comparable descending gradients, with photogrammetry
estimating a lower degree of softening resulting in a higher fracture energy and slip.
Although this difference in post peak slip is not shown in figure 4.12 (b), the strain
behaviour after the onset of debonding can be difficult to measure using
photogrammetry. The photogrammetric measurements generally require filtering to
smooth out irregularity and noise in the raw data. This irregularity can increase after the
onset of debonding occurs. It is therefore recommended that where a photogrammetry
data acquisition system is used, it always be verified with strain gauge data for
experiments of this nature.
Bond slip correlations for Anchorage type 1 specimens demonstrate a similar trend of
bond-slip curves for both photogrammetry and strain gauge data. It can be verified that
curves are very similar in terms of peak shear stress and corresponding slip.
4.4.6 Effective strain in FRP laminates used in design
The effective strain of a FRP laminate is a governing factor in the design of FRP
systems. It is the maximum strain that can be achieved in the FRP system prior to
failure considering all possible failure modes. (ACI 440.2R-02 2002) states that for
bonded U-wraps or bonded face plies; the maximum strain in the FRP which may be
used in design can be expressed in the following equation:
(4.4)
Where kv is a bond reduction coefficient applicable to shear and fu is the ultimate strain
of the FRP plate at rupture. The maximum strain in the FRP shown in equation 4.4
should be smaller than 0.004 for shear strengthened members only to ensure aggregate
interlock.
When this is computed using current specimen material properties (control specimen)
the effective FRP strain fe is calculated as: 0.00189 which is 75% of the observed
maximum elongation recorded. This value is considered acceptable for use in design
Chapter 4 – Experimental Investigation into FRP Anchorage Systems utilising a
Mechanically Strengthened Substrate
76
with a factor of safety of 0.75. Table 4.8 compares the effective FRP design strain with
the maximum elongations recorded in experimental testing. No prediction model
currently exists in literature for the anchorage solution which was tested.
Table 4.7 - Maximum FRP elongations and corresponding effective FRP strains and utilisation percentiles
4.5 Summary
This chapter presented the results of anchorage bond strength of FRP laminates bonded
to concrete with modified substrate conditions. The development of an effective
anchorage solution for the improvement of overall substrate and bond strength
properties has been presented using an experimental study. The results and discussions
presented allow the following conclusions to be made:
Design guidelines can limit the FRP material strains to levels as low as 10-25% of
the ultimate material strain at rupture.
By anchoring the ends of FRP laminates or sheets it is possible to achieve higher
bond strength capacities and degrees of utilisation of the FRP material resulting in
increased economy.
The strength of the concrete substrate is a key factor affecting the delamination mode
and overall bond strength.
The introduction of a mechanical chase cut into the concrete over the anchorage
length is an effective way to improve the strength of the concrete substrate, resulting
in higher FRP maximum elongations, bond stress, slip and load carrying capacities
Specimen Max
Elongation ( )
Effective FRP strain, fe
ACI440.2R-08, ( )
FRP Rupture strain,
( ) Fiber
Utilisation GA AR Control Specimen WG9 2535 2706 1890 14,000 18-19%
Stage 1 - Anchorage
WG1 4640 4434 NA 14,000 31-33%
WG2 4881 4733 NA 14,000 33-35%
Chapter 4 – Experimental Investigation into FRP Anchorage Systems utilising a
Mechanically Strengthened Substrate
77
The effect of the chase is a 95-100% increase in ultimate capacity, 118% increase in
bond stress and 83-93% increase in the maximum strain level reached prior to
failure.
3D non-contact measuring technique based on image correlation photogrammetry is
a viable measurement technology which results in good correlation with
measurements obtained using discrete strain gauges.
Chapter 5 – Experimental Investigation into FRP Anchorage System Utilising
Unidirectional and Bidirectional fiber Patch Anchors
78
5 CHAPTER 5 - EXPERIMENTAL INVESTIGATION INTO FRP ANCHORAGE SYSTEMS UTILISING UNIDIRECTIONAL AND BI-BIRECTIONAL FIBER PATCH ANCHORS
5.1 Introduction
The complex geometry of box girder bridges and together with the requirements for
shear and torsional retrofit pose many unique challenges in the area of FRP anchorage
and termination detailing where a lack of guidance currently exists in literature. The
development of FRP forces around the corners of the box section and sufficient
anchorage of the external FRP outer web reinforcement into the bridge deck where two
critical components in the FRP design termination detailing which required foremost
attention.
The appropriate anchorage of the externally bonded fibers into the deck and around the
corners of the box section should meet the criteria of providing the necessary continuity
to engage the torsional tensile stresses in addition to the adequate transfer of shear
forces into the bridge deck. The heavy strengthening demands and corresponding
material requirements posed additional anchorage criteria: to improve the FRP design
strains that may be safely adopted in design resulting and material and cost savings.
With the above criteria in mind a total of six anchorage schemes were devised and
investigated through laboratory testing program.
5.2 Specimen Design
The following anchorage solutions were designed to improve the efficiency of FRP
laminates applied to the sides of concrete webs for shear and torsional retrofit. Figure
5.1 highlights the intended practical application of the anchorage types investigated.
Anchorage types 4-6 were developed for applications to the web flange interfaces in an
attempt to improve the anchorage strength of FRP laminates applied to the outside faces
of the box section. The nature of the bidirectional fabric meant that it could also be
applied around the joint between the outer webs and the deck soffit of the box section.
Chapter 5 – Experimental Investigation into FRP Anchorage System Utilising
Unidirectional and Bidirectional fiber Patch Anchors
79
The later served dual purposes: to anchor the FRP laminates applied the outer webs and
to develop torsional hoop stresses between laminates applied to the outer webs and deck
soffit.
Figure 5.1 - Anchorage types 2 -5 applied to a box girder bridge.
The experimental program utilised two differing concrete prism dimensions suitable for
each anchorage type. Types 0, 2 and 4 utilised 2 no. (type A) reinforced concrete
blocks of dimension 250mm x 300mm x 600mm (figure 5.2 and 5.7). Types 3, 5 and 6
utilised 2 no. (type B) reinforced concrete blocks of dimension 200mm x 400mm x
600mm with a curved end recessed from the base of the prism (Figure 5.5, 5.9 and
5.12). Each face of the concrete block was reinforced nominally with 16mm diameter
bars (grade 500 MPa) spaced at 200mm centres to replicate the existing reinforcement
present in the box girder webs.
All specimens consisted of a single 120mmx2mmx1000mm laminate strip bonded to the
surface of the concrete block with a bond length of 500mm for concrete block type A
and 425mm for block type B. The slight difference in bond length between each type is
deemed acceptable due to the concept of effective bond length. Current bond strength
models predict a bond length beyond which any increase in the bond length cannot
increase the anchorage strength. This has been confirmed by many experimental studies
(Chajes et al. 1996; Maeda et al. 1997; Chen and Teng 2001); and is usually no greater
than 300mm.
Chapter 5 – Experimental Investigation into FRP Anchorage System Utilising
Unidirectional and Bidirectional fiber Patch Anchors
80
5.3 Test Preparation and Material properties
The face of the concrete blocks were cleaned and sandblasted to achieve a profile
similar to 60 grit sand paper prior to further surface preparation and application of the
laminate strip. In all cases, the first layer of fabric and its component saturant were
applied (wet-lay-up) and allowed sufficient time to reach a tacky state. This was
identified as the level of curing where by the saturant had set enough to hold the fabric
in place and also not be contaminated with laminate adhesive (approximately 30-45
minutes curing).
Application was followed by curing which occurred in a temperature controlled
chamber of 50°C for a period of 48 hrs, then further curing at room temperature (22°C)
for a further 72 hours prior to testing. The accelerated curing was necessary due to the
tight time constraints imposed by the experimental program and is not a requirement for
the performance of the system in field or laboratory conditions. Adhesion tests were
carried out on additional adhesion samples to verify laminate adhesion and curing prior
to testing. A summary of the test specimens constructed for anchorages types 2-6 are
summarized in table 5.1.
Anchorage Type Ref Anchorage Material No. of
plies Orientation
0 WG9 Contol na na 2 WG3 Unidirectional fiber 2 90 WG4 Unidirectional fiber 2 90 3 WG5 Unidirectional fiber 2 180 WG6 Unidirectional fiber 2 180 WG7 Unidirectional fiber 2 180 4 WG12 Bidirectional Fiber 1 ±45°5 WG10 Bidirectional Fiber 2 ±45° WG11 Bidirectional Fiber 2 ±45°
6 WG8 Unidirectional + Bidirectional Fibers 2+1 180/±45°
Table 5.1 – Summary of test specimens constructed in experimental program
Chapter 5 – Experimental Investigation into FRP Anchorage System Utilising
Unidirectional and Bidirectional fiber Patch Anchors
81
5.3.1 Anchor Type 2:
Anchorage type “type 2” is applicable for FRP anchorage at the web flange interfaces
(refer figure 5.1). This solution can also be applied to the webs of rectangular concrete
T-beams. The method comprised of using 2 plies of 250mm wide unidirectional FRP
fabric wrap (Mbrace CF140) applied horizontally across the laminate strip, as depicted
in figure 5.2. The direction of fabric fibers was 90° to the direction of laminate. The first
sheet overlayed the second, sandwiching the laminate strip in between. The anchorage
was developed in order to investigate the contribution of unidirectional fabric to resist
the tensile peeling stresses in the anchorage zone and to assess the potential for
distribution of fiber-adhesive stresses over a greater area of concrete.
Figure 5.2 - Anchorage types 2 specimen geometry and material properties
Chapter 5 – Experimental Investigation into FRP Anchorage System Utilising
Unidirectional and Bidirectional fiber Patch Anchors
82
(a) (b) (d)
Figure 5.3 - Construction process of Type 2 Anchorage Specimen; (a) Placement and rolling out of voids of the first layer of MBRACE CF140, positioned 90° to the direction of the laminate strip; (b) Profiling of laminate adhesive (as per manufacturers specification) to the surface concrete block over MBRACE CF140 fabric and application of application of MBRACE saturant; (c) Placement of second (top) layer of MBRACE CF140 sheet to concrete block directly over location of first layer.
5.3.2 Anchor Type 3:
Type 3 specimens utilised an anchorage consisting of 2 plies of unidirectional fibers
orientated parallel to the direction of the laminate. This detail was developed for use
where combined shear and torsional strengthening is a requirement. Full wrapping of
the concrete section is usually required for torsional strengthening and can be achieved
by using a continuous sheet of FRP fabric applied to all sides of the section; or in the
form of a U-wrap with appropriate anchorage into the flanges. Where the use of FRP
laminate ligatures in place of FRP fabric is preferred due to strength, economy or
practical requirements, a suitable detail to transfer the tensile forces around the section
corners is required in order to develop the torsional hoop stresses. Type 3 investigates
the application of L-shaped lengths of FRP unidirectional fabric to the corners of a box
section. These are appropriately lapped with a FRP laminate which is applied to the
main faces of the concrete prism as shown in figure 5.4. Three specimens were
constructed for type 3 (WG5, WG6, WG7) with a “dry” method of application used for
the last specimen (WG7). The alternative application procedure ensured that the
interface between each layer of FRP material had hardened sufficiently to ensure a
Chapter 5 – Experimental Investigation into FRP Anchorage System Utilising
Unidirectional and Bidirectional fiber Patch Anchors
83
“dry” joint had occurred. This “cold” formed method was used to replicate possible
work conditions/sequences on site. The specimen geometry and construction process for
anchor type 3 are depicted in figures 5.5 and 5.6.
Figure 5.4 - Anchorage types 2 and 3 applied to a box girder bridge.
Figure 5.5 - Anchorage types 3 specimen geometry and material properties.
Chapter 5 – Experimental Investigation into FRP Anchorage System Utilising
Unidirectional and Bidirectional fiber Patch Anchors
84
(a) (b) (c)
Figure 5.6 - Construction process of Type 3 Anchorage Specimen; (a) Rolling out voids of the first layer of CF140 sheet once applied to the concrete block; (b) Applying MBRACE laminate strip to prepared surface of concrete block; (e) Applying, rolling out and removing voids from between the laminate strip and second layer of CF140.
5.3.3 Anchor Type 4:
Anchor type 4 consisted of the application of a single 2mm thick x 120mm wide FRP
laminate to the concrete surface followed by the placement of 1 layer of bidirectional
fabric (270 mm wide) across the laminate. The fabric was wrapped around the corners
of the concrete and bonded a length of 50mm down the sides of the block (refer figure
5.7) with a fiber orientation that was ±45º to the direction of loading. The same
orientation of bidirectional fabric was also used for specimens that followed later in the
program (anchor types 5 and 6). The construction process for anchor type 4 is
summarised in figure 5.8.
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Unidirectional and Bidirectional fiber Patch Anchors
85
Figure 5.7 - Anchorage type 4 specimen geometry and material properties (WG12)
(a) (b)
Figure 5.8 - Construction process of Type 4 Anchorage Specimen; (a) Profiling and placement of laminate and adhesive (as per manufacturers specification) to the surface of the concrete block; (b) Placing and rolling out voids of TYFO BCC ±45° sheet, ensuring the direction of fibers is correct.
5.3.4 Anchor Type 5: Anchor type 5 utilised 2 layers of 270mm wide bidirectional fabric applied to the
concrete prism shown in figure 5.9. The first layer of fabric was initially bonded the
concrete prior to application of the laminate and was followed by the second fabric
layer, sandwiching the laminate in between. This is highlighted by examining the
construction process presented in figure 5.10. The fabric (400mm in length) was applied
to the top of the prism only with no side bonding being used.
Chapter 5 – Experimental Investigation into FRP Anchorage System Utilising
Unidirectional and Bidirectional fiber Patch Anchors
86
Figure 5.9 - Anchorage type 5 specimen geometry and material properties (WG10 & WG11)
(a) (b)
(c) (d)
Figure 5.10 - Construction process of Type 5 Anchorage Specimen; (a) Rolling out voids of in bidirectional fabric once applied to concrete block; (b) Applied laminate adhesive (as per manufacturers’ specification) to the surface of the bidirectional fabric and concrete block; (c) Laminate strip ready for application of top bidirectional fabric layer; (d) Completed anchorage specimen with two layers of TYFO BCC ±45° bidirectional fabric sheet, positioned ±45° to the direction of the laminate strip with laminate strip sandwiched in between.
Chapter 5 – Experimental Investigation into FRP Anchorage System Utilising
Unidirectional and Bidirectional fiber Patch Anchors
87
5.3.5 Anchor Type 6:
The combination of unidirectional and bidirectional fiber patch anchors was conceived
to address concrete box girder web-soffit transitions, where a continuity of fiber stresses
are required around the bend. As a result, the unidirectional fibers provide the
longitudinal trass transfer, whereas the bidirectional fibers provide the mechanism of
stress transfer between the FRP laminate and a wider width of unidirectional fibers,
hence facilitating a more efficient stress transfer. The intended application of the
anchorage to a box girder section is shown in figure 5.11. The detail was replicated in
anchor type 6, which consisted of 2 layers of unidirectional fabric applied to the
concrete, with a fiber orientation parallel to the laminate and subsequent direction of
loading. Following the application of both layers of unidirectional fabric (laminate
placed in between each layer), a single layer of bidirectional fabric (270mm
widex400mm long) was applied closest to the edge of loading without side bonding as
depicted in figure 5.12 and 5.13.
Figure 5.11 - Application of anchorage type 6 to proposed box girder section
Chapter 5 – Experimental Investigation into FRP Anchorage System Utilising
Unidirectional and Bidirectional fiber Patch Anchors
88
Figure 5.12 - Anchorage type 6 specimen geometry and material properties (WG8)
(b) (b)
Figure 5.13 - Construction process of Type 6 Anchorage Specimen; The construction sequence used for the Type 2 specimen used the following additional steps (a) Sand back surface of cured CF140 sheet (top sheet). (b) Placing and rolling out voids of TYFO BCC ±45° sheet, ensuring the direction of fibers is correct.
5.4 Experimental Results
5.4.1 Failure modes
The following section summarises the failure modes observed during the testing for the
various specimens anchored with unidirectional and bidirectional fabric.
Chapter 5 – Experimental Investigation into FRP Anchorage System Utilising
Unidirectional and Bidirectional fiber Patch Anchors
89
Anchor Type 2 - Two stages of delamination prior to failure were observed for
anchorage type 2. The first stage comprised of cover separation failure in the initial
225mm length of un-anchored laminate (refer figure 5.14). This was verified by a
sudden increase in strain in gauges G3 and G4 at a load level between 110-120 kN,
which indicated that debonding had progressed into the anchored portion of the
laminate. Cover separation failure occurred over a width which was greater than the
width of the laminate, which is apparent by the exposed concrete aggregate observed in
figure 5.14 either side of the laminate. This was attributed to the 50mm adhesive tappers
of applied to the edges of the. The tapper was applied throughout the full bonded length
of the laminate to provide a smooth transition for the unidirectional fabric wrap applied
horizontally across the laminate strip. The results suggest that the use of adhesive
tappers can effectively distribute stresses from the FRP laminate through the adhesive,
to a greater width of concrete and can potentially result in higher load carrying
capacities; however the extent remains to be quantified. Furthermore, with increasing
load application and additional partial debonding, the horizontal fibers of the FRP fabric
wrap were observed to incline in angle toward the direction of loading. The inclination
resulted in higher fabric strains, in addition to the strain induced by the tensile peeling
stresses between the laminate and the concrete. Failure in the anchored portion of the
laminate occurred within the concrete adhesive interface. The above failure mode was
consistent between both specimens tested.
Chapter 5 – Experimental Investigation into FRP Anchorage System Utilising
Unidirectional and Bidirectional fiber Patch Anchors
90
(a) (b) (c)
Figure 5.14 - Testing of type 2 (WGB3); (a) specimen ready for testing; (b) concrete rupture at adhesive - concrete interface; (c) shear rupture of CF140 fabric at point of wrap around;
Anchor Type 3 – An abrupt multi-phase failure was observed just prior to ultimate load
being reached. Debonding was first initiated between the concrete and adhesive which
was bonded to the first layer of FRP fabric. This was followed by debonding of the FRP
laminate from between both layers of FRP fabric. Failure within the concrete was
observed, localised to the width of the FRP laminate with vertical splitting of the fabric
occurring at the laminate edges, as shown in figure 5.15.
Chapter 5 – Experimental Investigation into FRP Anchorage System Utilising
Unidirectional and Bidirectional fiber Patch Anchors
91
(a) (b) (c)
Figure 5.15 - Testing of anchor type 3 (WG6) (a) specimen ready for testing; (b) Laminate bond failure at 1st and 2nd fabric layer interfaces; (c) 2nd layer of fabric rupture at base of laminate strip; (d) side view of debonded laminate strip from concrete block.
Anchor Type 4 – The following describes the first series of specimens which were
anchored using bidirectional fabric and exhibited multiple stages of delamination prior
to ultimate failure. Initially, debonding of the laminate/sheet to concrete interface
occurred at the loading edge and was followed by a combination of laminate debonding,
laminate rupture (along the direction of the fibers) and ±45° bidirectional fabric sheet
rupture (along the direction of the fibers). The stages of debonding for anchor type 4,
are summarised in figure 5.16.
Chapter 5 – Experimental Investigation into FRP Anchorage System Utilising
Unidirectional and Bidirectional fiber Patch Anchors
92
(a) (b) (c)
Figure 5.16 - Testing of anchor type 4 (WG12) (a) specimen ready for testing; (b) partial concrete-adhesive separation failure and fabric rupture (c) fabric rupture along the ±45° fiber direction.
Anchor Type 5 – Figure 5.17 highlights the multi-phase failure of both type 5 specimens
(WG10 and WG11) observed during testing. The first stage of concrete-adhesive
interfacial debonding of the laminate occurred in the initial 50mm of unanchored length
for both specimens. Specimen WG10 went on to show progressive debonding of the
sandwiched laminate structure from the concrete surface, which resulted in complete
debonding of the laminate and bidirectional fabric structure from the concrete block. It
is believed that this mode of failure was induced by an inadequate surface roughness,
caused by the recycling of the (type B) concrete blocks (used in type 3 specimens) and
the need for secondary sand blasting to remove existing bonded fabric. The remaining
stages of delamination for specimen WG11 were a combination of laminate debonding,
laminate rupture (along the direction of the fibers) and ±45° bidirectional fabric sheet
rupture (along the direction of the fibers).
Chapter 5 – Experimental Investigation into FRP Anchorage System Utilising
Unidirectional and Bidirectional fiber Patch Anchors
93
(a) (b) (c)
Figure 5.17 - Testing of anchor type 5 (WG10); (a) specimen ready for testing; (b) and (c) delamination of sandwiched laminate at adhesive-concrete interface.
Anchor Type 6 – The combination of unidirectional and bidirectional fabric, used in
anchor type 6, significantly enhanced the anchorage strength of the specimen. The
system remained in-tact (without signs of debonding) until rupture of the FRP laminate,
which is depicted in figure 5.18.
(a) (b) (c)
Figure 5.18 - Testing of anchor type 6 (WG8) (a) specimen ready for testing; (b) and (c) ruptured laminate (parallel to fiber direction); (c) close up of laminate failure over specimen free length;
Chapter 5 – Experimental Investigation into FRP Anchorage System Utilising
Unidirectional and Bidirectional fiber Patch Anchors
94
5.4.2 FRP strain distributions along length of laminate
Table 5.2 summarises the failure loads and maximum FRP elongations reached in types
0, 2-6 of the experimental program. In tables and figures which follow reference is
made to AR (Photogrammetry) and SG (strain gauge). These refer to the two data
acquisition techniques used in the experimental programme.
Table 5.2 – Maximum FRP elongations and corresponding effective FRP strains and utilisation percentiles (types 0, 2-6)
(a) (b)
0
500
1000
1500
2000
2500
3000
0 100 200 300 400
Micro
strain
()
Distance From Gauge G1 (mm)
80kN (SG)
80kN(AR)
96.6kN (SG)
96.6kN(AR)
40kN (SG)
40kN(AR)0
500
1000
1500
2000
2500
3000
3500
0 100 200 300 400
Micro
strain
()
Distance From Gauge G1 (mm)
60kN(AR)
100kN(AR)
137.3kN(AR)
60kN (SG)
100kN (SG)
137.3kN (SG)
Type Ref Pmax
Max
Laminate
strain ( )
Incre
-se in
Load
Max strain in
FRP ±45°
Fabric (SG)
Max strain in
FRP ±45°
Fabric (AR)
Failure
Mode
GA/AR (kN)
GA ( )
AR ( ) %
LS ( )
RS ( )
LS ( )
RS ( )
0 WG9 99.6 2535 2706 - - - - - CSF
2 WG3 138.2 3242 3212 27.9 - - - - CSF WG4 142 3142 3235 23.9 - - - - CSF/ ASF
3 WG5 156.5 3470 3607 36.9 - - - - CSF/ ASF WG6 146 3239 3488 27.8 - - - - CSF/ ASF WG7 145.3 3245 3204 28 - - - - CSF/ ASF
4 WG12 218.3 5800 4867 128.8 12896 13632 13136 - CSF /
PLR/ PFR
5 WG10 213 4900 5261 93.3 5228 5225 3982 - CSF WG11 236.9 5300 - 109.1 7433 12834 - - CSF
6 WG8 261.4 7500 7589 195.9 4177 4372 4054 - PASF/LR Note: CFS (Cover separation failure); ASF (Adhesive separation failure); PASF (Partial Adhesive
Separation Failure); LR (Laminate Rupture); FR (Fabric Rupture); LS (fabric right ride of laminate); RS (fabric right side of laminate)
Chapter 5 – Experimental Investigation into FRP Anchorage System Utilising
Unidirectional and Bidirectional fiber Patch Anchors
95
(c) (d)
(e) (f)
(g) (h)
0
500
1000
1500
2000
2500
3000
3500
0 100 200 300 400 500
Micro
strain
()
Distance From Gauge G1 (mm)
40kN (SG)
40kN(AR)
100kN (SG)
100kN(AR)
137kN (SG)
137kN(AR)
0
500
1000
1500
2000
2500
3000
3500
4000
0 50 100 150 200 250 300
Micro
strain
()
Distance From Gauge G1 (mm)
60kN(AR)100kN(AR)156.4kN(AR)60kN (SG)100kN (SG)156.4kN (SG)
0
500
1000
1500
2000
2500
3000
3500
0 100 200 300 400
Micro
strain
()
Distance From Gauge G1 (mm)
40kN (SG)40kN(AR)100kN (SG)100kN(AR)142.8kN (SG)142.8kN(AR)
0
500
1000
1500
2000
2500
3000
3500
0 100 200 300 400
Micro
strain(
)
Distance From Gauge G1 (mm)
60kN(AR)
110kN(AR)
143.8kN(AR)
60kN (SG)
110kN (SG)
143.8kN (SG)
0
1000
2000
3000
4000
5000
6000
0 100 200 300 400 500 600
Micro
strain
()
Distance From Gauge G1 (mm)
125kN (SG)125kN(AR)175kN (SG)175kN(AR)217.2kN (SG)
0
1000
2000
3000
4000
5000
6000
0 100 200 300 400
Micro
strain
()
Distance From Gauge G1 (mm)
125kN (SG)125kN(AR)175kN (SG)175kN(AR)211.95kN (SG)211.95kN(AR)
Chapter 5 – Experimental Investigation into FRP Anchorage System Utilising
Unidirectional and Bidirectional fiber Patch Anchors
96
(i) (j)
Figure 5.19 - Strain vs distance along Laminate; (a) Type 0 (Control) ; (b) Anchorage Type 2 (WG3); (c) Anchorage Type 2 (WG4); (d) Anchorage Type 3 (WG5); (e) Anchorage Type 3 (WG6); (f) Anchorage Type 3 (WG7); (g) Anchorage Type 4 (WG12); (h) Anchorage Type 5 (WG10); (i) Anchorage Type 5 (WG11); (j) Anchorage Type 6 (WG8);
FRP elongation along the length of the laminate are reported in figures 5.19 for
anchorage types 0 and 2-6. An examination of the experimental data shows that anchor
type 2 was effective in increasing the ultimate failure load by 39-43% and resulted in an
increase in the maximum laminate strain of 19-28% prior to failure. The higher load
carrying capacity of the anchorage was mainly attributed to the 50mm adhesive tapers
distributing the laminate-adhesive stresses to a greater width of concrete and the
addition of the unidirectional fabric contributing to resist load through a strut-tie action
resulting from the fabric fibers inclining towards the direction of loading prior to
failure. Close correlations are observed between the photogrammetry and strain gauge
measurements. The deviations in strain seen in figure 5.19(b) at location (300mm) prior
to failure are due to the strain gauge G7 slipping. Photogrammetry data showed a
continuous strain profile along the length of the laminate at each load increment. As a
result, a slight dip in strain level was revealed at a location of 50mm from gauge G1,
which corresponded to the edge of the concrete block (refer figure 5.19 (a), (b), (d) and
(f).
The utilisation of unidirectional fabric applied parallel to the direction of the laminate
(anchorage type 3) was effective in increasing the ultimate failure load by 46-57%
compared to the unanchored control specimen. An increase in maximum laminate
elongation of 18-37% was attributed to this form of anchorage. The increase in
0
500
1000
1500
2000
2500
3000
3500
4000
4500
0 50 100 150 200 250 300
Micro
strain(
)
Distance From Gauge G1 (mm)
75kN (SG)
100kN (SG)
125kN (SG)
150kN (SG)
175kN (SG)
180kN (SG)
0
1000
2000
3000
4000
5000
6000
7000
8000
0 100 200 300 400
Micro
strain
()
Distance From Gauge G1 (mm)
125kN (SG)125kN(AR)175kN (SG)175kN(AR)260.8kN (SG)260.8kN(AR)
Chapter 5 – Experimental Investigation into FRP Anchorage System Utilising
Unidirectional and Bidirectional fiber Patch Anchors
97
anchorage strength observed in table 5.2 was due to the transfer of bond stress to a
greater distance away from the loaded edge, resulting in an increased effective
anchorage length. This is clearly observed in figure 5.19(e) by the higher level of strain
recorded at a distance of 300mm away from the loaded face prior to failure. The
presence of micro-cracking at the concrete-to-adhesive interface and consequent slip of
the laminate during loading gradually forced a re-distribution of bond-stress further
away from the loaded face. It is believed that this redistribution was greatly facilitated
by the anchoring effect of the unidirectional fabric curved and anchored around the end
of the concrete block.
The use of bidirectional FRP fabric to anchor the FRP laminate was adopted in
anchorage types 4 and 5. It is noted that an increase in failure load of 128% was
observed for the type 4 anchor, due to the application of one ply of bidirectional fabric
anchored 50mm down the sides of the concrete block. A 93-109% increase in failure
load was reached in type 5 which utilised two plies with no fabric anchorage. The
maximum ±45° fabric elongations measured suggest that the fabric strain utilisations
were 2-3 times greater when using a single fabric ply with anchorage compared with
where no anchorage was provided. The 50mm anchorage was omitted in type 5 where
the 2 plies of ±45° fabric were bonded across the full width of the type B concrete
block.
Anchorage type 6 showed the greatest increase in ultimate failure load (195%) and
failed by rupture of the FRP laminate at a load of 261.4 KN. The maximum ±45° fabric
strain reached was 3762-4054 with a corresponding laminate strain of approximately
7500 . Examining the strain distribution prior to failure in figure 5.19 (j) shows the
specimen to have the highest effective anchorage length where laminate strains of 3458
were recorded at a distance of 400mm away from the loaded face. By introducing the
±45° bidirectional fabric sheet in addition to the unidirectional fibers installed parallel
to the direction of the laminate, the anchorage has combined the benefits of anchor
types 3 and 5, which resulted in a distribution of fiber-to-adhesive bond stresses over a
greater length and width of concrete. The strains in the ±45° bidirectional fabric reached
similar levels to those recorded in type 5, 3762-4054 which indicated a comparable
distribution of stress across the width of the concrete block between the two samples.
Chapter 5 – Experimental Investigation into FRP Anchorage System Utilising
Unidirectional and Bidirectional fiber Patch Anchors
98
5.4.3 Load – Displacement curves
The stages of delamination can be observed from the load displacement curves
presented in figures 5.20 to 5.23. The onset of de-bonding of the FRP-to-concrete
interface occurs with an increase in transverse micro cracking and a local reduction of
bond stiffness which results in a flattening of the load-displacement curve and a
redistribution of strain and consequent increase in the strain levels in areas further away
from the loaded edge. This can be clearly observed in all specimens.
Anchor type 2 – Examining the strain distributions from figures 5.20, three stages of
debonding are observed in two zones along the length of the laminate. The figure
highlights the first stage of delamination commencing between 40-80kN. This
corresponds to gauges G3 and G4 which are located 50-100mm away from the loaded
face. No appreciable strains were recorded by gauges G5 and G6 (located 150-200mm
from the loading edge of the concrete block) until a load of approximately 110-120kN
was reached. Complete debonding of the laminate occurred two stages. The un-
anchored 225mm length of bonded laminate failed first at load levels in the order of
120kN. A further load increase of 20-30 kN occurred after the first stage of laminate
debonding had occurred due to the restraint provided by the FRP fabric anchorage.
Close correlations are observed between the photogrammetry and strain gauge
measurements. The deviation in strain seen in figure 5.20 (b) at location G7 prior to
failure is due to the strain gauge G7 slipping. The photogrammetry data showed a
continuous strain profile along the length of the laminate at each load increment.
Examining the photogrammetry data, a slight dip in strain level is observed at a location
of 50mm. This has been consistently observed in specimens WG3, WG6 and WG7. The
location of strain depression corresponds to the edge of the concrete block.
Anchor type 3 – Strain distribution along the length of the laminate, presented in figure
5.21, shows three significant regions. Initial bond failure commenced within the initial
100mm between a load of 50kN and 100kN. This is clearly observed by the reduced
slope of the load-strain curve within this zone. Further debonding at a distance of
200mm from the loading edge commenced after loads exceeded 120kN-140kN. Test
Chapter 5 – Experimental Investigation into FRP Anchorage System Utilising
Unidirectional and Bidirectional fiber Patch Anchors
99
specimens WG5-WG7 showed significant strain increases of gauges G3-G6 at this load
level. Ultimate debonding and specimen failure occurred at ultimate load. The two data
acquisition techniques showed good correlations throughout all loading increments.
Tables 5.2 and 5.3 illustrate a similar trend of failure load and maximum bond stress
across specimens WG5-7 which illustrates that the “dry” application method did not
significantly affect the ultimate strength and failure strain of the specimens.
The increase in anchorage strength observed in table 5.2 was attributed to a transfer of
bond stress to a greater distance away from the loaded edge. This is clearly observed in
figure 5.21 by the higher level of strain recorded at a distance of 300mm away from the
loaded face. This resulted in an increased effective anchorage length.
Anchor type 4 – The first series of anchorages which utilised bidirectional FRP fabric to
anchor the FRP laminate was adopted in anchorage types 4. The application of one ply
of bidirectional fabric anchored 50mm down the sides of the concrete block resulted in a
maximum laminate strain of 5800 and an increase in failure load of 128% (figure
5.22). Photogrammetry data was available up to a load level of 175 kN beyond which
recording was interrupted.
Anchor type 5 – A 93-109% increase in failure load and maximum strain level of 4900-
5300 were reached in specimen type 5 which utilised two plies of bidirectional fabric
applied to the loaded face only. The increase is somewhat lower than that reached in
type 4 which suggests that the 50mm development of bidirectional fabric down the sides
of the concrete block used in type 4 significantly increased the effective utilisation of
the fabric. The maximum ±45° fabric elongations measured suggest that the fabric strain
utilisations were 2-3 times greater when using a single fabric ply with anchorage
compared with where no anchorage was provided. The load-displacement distributions
are presented in figure 5.23.
Anchor type 6 – The specimen failed through a combination of partial adhesive
separation failure and rupture of the FRP laminate at a load of 261.4 KN and showed
the greatest increase in ultimate failure load (195%) with a corresponding laminate
strain of approximately 7500 . The maximum ±45° fabric strain reached was 4372 -
Chapter 5 – Experimental Investigation into FRP Anchorage System Utilising
Unidirectional and Bidirectional fiber Patch Anchors
100
4054 . The strains in the ±45° bidirectional fabric reached similar levels to those
recorded in type 5, which indicated a comparable distribution of stress across the width
of the concrete block between the two samples. Examining the strain distribution prior
to failure in figure 5d demonstrates the specimen to have the highest effective
anchorage length where laminate strains of 3458 were recorded at a distance of
400mm away from the loaded face. By introducing the ±45° bidirectional fabric sheet in
addition to the unidirectional fibers installed parallel to the direction of the laminate; the
anchorage has combined the benefits of anchor types 3 and 5, resulting in a distribution
of fiber-to-adhesive bond stresses over a greater length and width of concrete.
An examination of the load-strain curves in figure 5.24 can provide some insight into
the stages of delamination during loading. Typically, the onset of de-bonding of the
FRP-to-concrete interface occurs together with an increase in transverse micro cracking.
This is followed by a local reduction of bond stiffness which results in a flattening of
the load-displacement curve and a redistribution of strain in the laminate to areas further
away from the loaded edge. Large increases in strain levels at a given load are indicative
of sudden debond and is usually followed by corresponding sudden strain increases in
gauges toward the end of the laminate. The introduction of the bidirectional fabric has
resulted in a gradual progressive delamination with increasing loading when compared
to the control specimen, which delaminated suddenly between 80 – 90 kN. This is
evidence of an improved ductility in the anchorage zone.
(a)
0102030405060708090
100110120130140
0 500 1000 1500 2000 2500 3000 3500
Load
(kN)
Micro strain ( )
G1 (SG)
G2 (SG)
G3 (SG)
G4 (SG)
G5 (SG)
G6 (SG)
G7 (SG)
G2 (AR)
G3 (AR)
G4 (AR)
G5 (AR)
G6 (AR)
G7 (AR)
Chapter 5 – Experimental Investigation into FRP Anchorage System Utilising
Unidirectional and Bidirectional fiber Patch Anchors
101
(b)
Figure 5.20 - Load vs strain distribution; (a) Anchorage Type 2 (WG3); (b) Anchorage Type 2 (WG4);
(a)
(b)
0102030405060708090
100110120130140150
0 500 1000 1500 2000 2500 3000 3500
Load
(kN)
Micro strain ( )
G1 (SG)
G2 (SG)
G3 (SG)
G4 (SG)
G5 (SG)
G6 (SG)
G7 (SG)
G2 (AR)
G3 (AR)
G4 (AR)
G5 (AR)
G6 (AR)
G7 (AR)
0102030405060708090
100110120130140150160
0 500 1000 1500 2000 2500 3000 3500 4000
Load
(kN)
Micro strain ( )
G1 (SG)
G2 (SG)
G3 (SG)
G4 (SG)
G5 (SG)
G6 (SG)
G7 (SG)
G3 (AR)
G4 (AR)
G5 (AR)
G6 (AR)
G7 (AR)
0102030405060708090
100110120130140150
0 500 1000 1500 2000 2500 3000 3500
Load
(kN)
Micro strain ( )
G1 (SG)
G2 (SG)
G3 (SG)
G4 (SG)
G5 (SG)
G6 (SG)
G7 (SG)
G3 (AR)
G4 (AR)
G5 (AR)
G6 (AR)
G7 (AR)
Chapter 5 – Experimental Investigation into FRP Anchorage System Utilising
Unidirectional and Bidirectional fiber Patch Anchors
102
(c) Figure 5.21 - Load vs strain distribution; (a) Anchorage Type 3 (WG5); (b) Anchorage Type 3 (WG6); Type 3 (WG7);
Figure 5.22 - Load vs strain distribution, Anchorage Type 4 (WG12);
(a)
0102030405060708090
100110120130140150
0 500 1000 1500 2000 2500 3000 3500
Load
(kN)
Micro strain ( )
G1 (SG)
G2 (SG)
G3 (SG)
G4 (SG)
G5 (SG)
G6 (SG)
G7 (SG)
G2 (AR)
G3 (AR)
G4 (AR)
G5 (AR)
G6 (AR)
G7 (AR)
0
2040
6080
100
120140
160180
200220
0 1000 2000 3000 4000 5000 6000
Load
(kN)
Micro strain ( )
G1 (SG)
G2 (SG)
G3 (SG)
G4 (SG)
G5 (SG)
G6 (SG)
G7 (SG)
G2 (AR)
G3 (AR)
G4 (AR)
G5 (AR)
G6 (AR)
G7 (AR)
020406080
100120140160180200220
0 1000 2000 3000 4000 5000 6000
Load
(kN)
Micro strain ( )
G1 (SG)
G2 (SG)
G3 (SG)
G4 (SG)
G5 (SG)
G6 (SG)
G7 (SG)
G2 (AR)
G3 (AR)
G4 (AR)
G5 (AR)
G6 (AR)
G7 (AR)
Chapter 5 – Experimental Investigation into FRP Anchorage System Utilising
Unidirectional and Bidirectional fiber Patch Anchors
103
(b)
Figure 5.23 - Load vs strain distribution; (a) Anchorage Type 5 (WG10); (b) Anchorage Type 5 (WG11);
(g)
Figure 5.24 - Load vs strain distribution; (a) Anchorage Type 6 (WG8);
5.4.4 Experimental bond slip curves
An understanding of the local bond–slip behaviour of the FRP-concrete interface is of
fundamental importance to the accurate modelling of debonding failures in FRP-
strengthened RC structures. The bond slip data can be computed from the axial strains
of the FRP plate measured at discrete locations. The strain measurements obtained at
gauge locations and from photogrammetry measurements can be used to obtain bond-
slip information. The bond slip relations are presented in equations 4.2 and 4.3. A
020406080
100120140160180200220240
0 1000 2000 3000 4000 5000 6000
Load
(kN)
Micro strain ( )
G1 (SG)
G2 (SG)
G3 (SG)
G4 (SG)
G5 (SG)
G6 (SG)
G7 (SG)
020406080
100120140160180200220240260280
0 1000 2000 3000 4000 5000 6000 7000 8000
Load
(kN)
Micro strain ( )
G1 (SG)
G2 (SG)
G3 (SG)
G4 (SG)
G5 (SG)
G6 (SG)
G7 (SG)
G2 (AR)
G3 (AR)
G4 (AR)
G5 (AR)
G6 (AR)
G7 (AR)
Chapter 5 – Experimental Investigation into FRP Anchorage System Utilising
Unidirectional and Bidirectional fiber Patch Anchors
104
comparison of the results obtained using both data acquisition techniques is presented in
figures 5.25 and 5.26. The shear stress of a particular location along the length of the
laminate can be found using a difference formula, while the corresponding slip can be
found by a numerical integration of the measured axial strains of the plate (Lu, Teng et
al. 2005). Table 5.3 summarises the maximum bond stresses and corresponding slips
measured as 50-75mm and 125mm away from the concrete free edge.
Table 5.3 – Bond stress and corresponding slip results summary (type 0, 2-6)
Type Ref Distance along laminate from Concrete free Edge 50-75mm 125mm
so so 0 WG9 SG 4.9 0.036 5.02 0.06
AR 3.25 0.06 5.83 0.11
2 WG3 SG 2.63 0.05 9.57 0 .16
AR 3.82 0.071 8.73 0.15 WG4 SG 3.94 0.087 4.35 0.06
AR 4.05 0.076 4.11 0.053
3
WG5 SG 3.26 0.07 4.59 0.083 AR 4.46 0.06 5.14 0.08
WG6 SG 7.94 0.36 5.34 0.14 AR 5.93 0.25 5.22 0.15
WG7 SG 6.16 0.30 5.37 0.10 AR 7.71 0.22 5.34 0.18
4 WG12 SG 6.99 0.19 7.73 0.15 AR 6.47 0.097 7.24 0.26
5
WG10 SG 14.64 0.87 7.18 0.30 AR 12.95 0.99 5.90 0.24 WG11 SG 15.94 0.43 7.32 0.17 AR - - - -
6 WG8 SG 15.46 1.92 7.54 0.11 AR 14.97 1.28 4.01 0.12
Chapter 5 – Experimental Investigation into FRP Anchorage System Utilising
Unidirectional and Bidirectional fiber Patch Anchors
105
(a) (b)
Figure 5.25 - Bond-slip curves fitted with Popovics equation at bond critical regions- (a) Type 0 (Control) ; (b) Anchorage Type 2 (WG4)
.
(a) (b)
(c) (d)
Figure 5.26 – Apparent Bond-slip curves fitted with Popovics equation at bond critical regions (measured 125mm away from Concrete free Edge) - (a) Anchorage Type 3 (WG6); (b) Anchorage Type 4 (WG12); (c) Anchorage Type 5 (WG10); (d) Anchorage Type 6 (WG8);
0
1
2
3
4
5
6
0 0.1 0.2 0.3 0.4
Bond
Stress
(MPa
)
Slip (mm)
Popovics
175mm(ARAMIS)175mm(GAUGE)
125mm
125 mm
0
1
2
3
4
5
0 0.1 0.2 0.3 0.4
Bond
Stress
(MPa
)
Slip (mm)
Popovics
175mm(ARAMIS)175mm(GAUGE)
125 mm
125 mm
0
1
2
3
4
5
6
0 0.1 0.2 0.3 0.4
Bond
Stress
(MPa
)
Slip (mm)
Popovics
175mm(ARAMIS)175mm(GAUGE)
125 mm
125 mm
0
1
2
3
4
5
6
7
8
0 0.1 0.2 0.3 0.4
Bond
Stress
(MPa
)
Slip (mm)
Popovics
175mm(ARAMIS)175mm(GAUGE)
125 mm
125 mm
0
1
2
3
4
5
6
7
8
0 0.2 0.4 0.6 0.8
Bond
Stress
(MPa
)
Slip (mm)
Popovics
175mm(ARAMIS)175mm(GAUGE)
125 mm
125 mm
0
1
2
3
4
5
6
7
8
0 0.1 0.2 0.3 0.4 0.5
Bond
Stress
(MPa
)
Slip (mm)
Popovics
175mm(GAUGE)125 mm
Chapter 5 – Experimental Investigation into FRP Anchorage System Utilising
Unidirectional and Bidirectional fiber Patch Anchors
106
A review of the bond-slip curves shows a comparable relationship between the two data
acquisition techniques. A distinction has been drawn between true and apparent bond
stress which is presented in figures 5.25 and 5.26. True bond stress can be defined as the
stress induced in the concrete as a result of a differential in strain measured across a
finite length along the FRP laminate. The true bond stress must be calculated from the
laminate strain and relies on the assumption of perfect strain compatibility between the
laminate, epoxy and concrete. Due to the presence on unidirectional and bidirectional
fabric layers for anchorage types 3-6, laminate strain readings were taken from
uppermost layer of FRP fabric. It is estimated that the strains measured from the
uppermost fabric layer will be different to the actual strain in the laminate. This is due
to the effects of interface slip between the fabric, laminate and concrete layers during
loading resulting in relaxation. In addition, shedding strains are induced in the ±45°
bidirectional fabric from the shedding of laminate forces to a wider area concrete, which
will not be felt by the FRP laminate. The true bond stress in the concrete for anchorage
types 3-6 is expected to be significantly lower. As a result, the bond stresses presented
have been defined as apparent stresses as a result of the strains measured from the
uppermost FRP fabric layer not corresponding to the true strain in the laminate and
concrete. A comparison of the bond-slip curves yields maximum bond stresses of 4.5-
5.5 MPa for both the control and anchorage type 3 & 4 specimens, the anchorage using
unidirectional fibers, was therefore un-successful in increasing the strength of the FRP-
concrete contact bond strength.
The softening branches of the bond slip curves follow comparable descending gradients
for anchorage specimens 2 and 5, with photogrammetry estimating a lower degree of
softening and a higher fracture energy and slip for specimens 0, 3 and 4. The difficulty
of obtaining accurate bond-slip curves is largely attributed to local variations in the
strains along the length of the laminate. This is clearly observed in figure 5.19 and was
due to the discrete nature of the concrete cracks, the heterogeneity of concrete and the
roughness of the underside of the debonded FRP plate (Teng et al. 2006). This variation
in strain was more pronounced in the photogrammetry measurements which required a
high degree of filtering to smooth out irregularity and noise in the raw data.
Chapter 5 – Experimental Investigation into FRP Anchorage System Utilising
Unidirectional and Bidirectional fiber Patch Anchors
107
It is noted that table 5.3 presents peak apparent bond stresses of up to 12.95-15.94 MPa
within the zone of 100-125mm from strain gauge G1 for anchorage types 5 and 6. This
zone corresponds to 50-75mm from the face of the concrete block. The stress slip
distribution within the zone demonstrated a linear trend with no indication of softening.
It is believed that the high level of apparent bond stress and the lack of softening are due
to the pronounced effects of interfacial slip between multiple fiber layers within this
zone. The effects of interfacial slip become less apparent at a distance of 175mm away
from strain gauge G1. The bond-slip curves within this zone indicate a softening tend
comparable with current prediction models. The presence of more two peaks in the
apparent bond stress distributions of figures 5.26 (b) and (c) are possibly related to the
presence of transverse concrete cracks which introduce local disturbances to the bond
behaviour.
5.4.5 Strain in bidirectional fibers
As depicted in figures, 5.27, 5.28 and 5.29 strain gauges were placed at certain intervals,
left and right of the laminate strip in order to capture the strains in the bidirectional
fibers. As a result, the orientation of the gauges was at ±45º parallel to the principle
direction of the fibers. It can be clearly observed that the strains in the bidirectional
fibers are generally maximum at, or near the laminate edge and dissipate to zero, over a
distance of approximately 60mm. Anchorage Type 4 experienced higher strains in the
bidirectional fibers (above 12000 ), since the sheet was anchored at right angles,
50mm down the sides of the concrete block. Where no 50mm tapers were used to
anchor the bidirectional sheet, a lower fiber strain of 3000-5000 was observed. The
distributions of strains away from the laminate edge provide useful information on the
effective length of the bidirectional fibers and subsequent width of the patch anchors,
which is the subject of investigation in stage 2 of the experimental program.
Chapter 5 – Experimental Investigation into FRP Anchorage System Utilising
Unidirectional and Bidirectional fiber Patch Anchors
108
Figure 5.27 – Strain of 45º Bidirectional FRP either side of laminate; Anchorage Type 4 (WG12)
Figure 5.28 – Strain of 45º Bidirectional FRP either side of laminate; Anchorage Type 5 (WG10)
0
2000
4000
6000
8000
10000
12000
14000
150 120 90 60 30 0 30 60 90 120 150
Microstrain
()
Distance from across concrete block from centre of laminate (mm)
90kN (SG L)
110kN (SG L)
130kN (SG L)
150kN (SG L)
170kN (SG L)
200kN (SG L)
90kN (SG R)
110kN (SG R)
130kN (SG R)
150kN (SG R)
170kN (SG R)
200kN (SG R)
0
1000
2000
3000
4000
5000
6000
150 120 90 60 30 0 30 60 90 120 150
Microstrain(
)
Distance from across concrete block from centre of laminate (mm)
50kN (SG L)
90kN (SG L)
130kN (SG L)
170kN (SG L)
200kN (SG L)
50kN (SG R)
90kN (SG R)
130kN (SG R)
170kN (SG R)
200kN (SG R)
Chapter 5 – Experimental Investigation into FRP Anchorage System Utilising
Unidirectional and Bidirectional fiber Patch Anchors
109
Figure 5.29 – Strain of 45º Bidirectional FRP either side of laminate; Anchorage Type 6 (WG8)
5.5 Summary
The experimental study was conducted to improve the efficiency and strain utilisations
of FRP bonded to concrete using unidirectional and bidirectional fabric anchorage
systems. The anchorages tested were successful in improving the degree of FRP strain
utilisation. The results and discussions presented allow the following conclusions to be
made:
Anchoring the ends of FRP laminates using unidirectional FRP fabric wrap applied
horizontally across the laminate strip (anchorage type 2) was effective in increasing
the ultimate failure load by 39-43% and resulted in an increase in the maximum
laminate strain of 19-28%.
The use of 50mm adhesive tappers increase along the length of the laminate was
found to distribute the laminate-adhesive stresses to a greater width of concrete.
FRP fabric applied horizontally across the laminate strip does not provide an
effective level of confinement to uniformly increase the bond strength between the
adhesive and concrete layer.
0
500
1000
1500
2000
2500
3000
3500
4000
150 120 90 60 30 0 30 60 90 120 150
Microstrain
()
Distance from across concrete block from centre of laminate (mm)
90kN (SG L)
110kN (SG L)
130kN (SG L)
150kN (SG L)
200kN (SG L)
240kN (SG L)
90kN (SG R)
110kN (SG R)
130kN (SG R)
150kN (SG R)
200kN (SG R)
240kN (SG R)
Chapter 5 – Experimental Investigation into FRP Anchorage System Utilising
Unidirectional and Bidirectional fiber Patch Anchors
110
The application of unidirectional fibers with an orientation parallel to the direction
of the laminate (anchorage type 3) was effective in increasing the ultimate failure
load by 46-57%. The overall increase in strength of this anchorage system was
attributed to the transfer of bond stress to a greater distance away from the loaded
edge, which was facilitated by the anchoring effect of the unidirectional fabric
curved and anchored around the end of the concrete block.
One ply of bidirectional fabric anchored 50mm down the sides of the concrete
block used to anchor the laminate in type 4 of the program was effective in
increasing the ultimate failure load by 128%.
The use of 2 plies of bidirectional fabric with no anchorage down the side of the
concrete block was effective in providing a 93-109% increase in failure load.
Bidirectional fabric applied to the ends of FRP laminates resulted in a more
efficient distribution of FRP-adhesive stresses over a greater width of concrete.
Utilising the properties of anchorage types 3 and 5 resulted in a distribution of
fiber-to-adhesive bond stresses over a greater length and width of concrete
achieving an increase in failure load of 195% and resulting in laminate rupture.
Chapter 6 – Experimental Investigation into the Size Effect of Bidirectional Fiber Patch
Anchors
111
6 CHAPTER 6 – EXPERIMENTAL INVESTIGATION INTO THE SIZE EFFECT OF BIDIRECTIONAL FIBER PATCH ANCHORS
6.1 Introduction
The first stage of the experimental results derived from the patch anchor specimens
showed very promising results. Of the six types of anchorage configurations
investigated in stage 1, anchor type 5, which used ±45º bidirectional fiber, was proven
to be the most efficient and versatile. As a result, it was decided that all future study
should focus exclusively on this anchor.
Since the stage 1 experiments were limited by case dependency and the relatively small
sample sizes employed. Many parameters remain to be investigated which could
influence the performance of the patch anchors when applied to structures containing
different material properties and design configurations. Factors such as: Concrete
strength, laminate thickness, laminate modulus and patch anchor size and their effect on
anchor performance remain to be quantified. Consequently, a further experimental study
was designed (herein stage 2) to investigate factors such as patch anchor size, laminate
thickness, laminate width and concrete strength.
6.2 Experimental Program
6.2.1 Specimen Design
The following stage of the experimental program (stage 2), consisted of patch anchor
configurations similar to those used in stage 1 - which were based on 2 plies of
bidirectional fabric, with the laminate sandwiched in between. However, the study was
designed to investigate a more commonly used laminate thickness (1.4mm) as opposed
to (2mm) which was used in stage 1. The laminate width adopted was also reduced from
120mm (stage 1) to 100mm in stage 2. The reduction in laminate width was chosen
specifically to observe the effect of laminate width and its relation to anchorage
strength.
Chapter 6 – Experimental Investigation into the Size Effect of Bidirectional Fiber Patch
Anchors
112
Another objective was to determine the size effect of the FRP patch anchor on the
overall anchorage strength. In shear strengthening applications, FRP laminates are often
installed to beams webs, side-by-side at a predefined spacing. Such situations require a
continuous form of anchorage applied to the FRP ends and the question naturally arises
regarding the relationship between laminate spacing and anchorage effectiveness.
Where continuous patch anchors are used, it is apparent that each laminate will transfer
bond stresses to a width of patch anchorage which is governed by the distance between
adjacent laminates (laminate spacing). In order to assess the performance of patch
anchorages under such situations, three alternative concrete block widths: 420, 320 and
220mm were chosen for further study.
Appropriate boundary conditions of symmetry at the concrete block left and right edges
were applied by replicating restraint normal to the concrete sides (x direction) whilst
allowing movement in the vertical plane (y direction). Such boundary conditions are
typically applied to replicate symmetry – in this case, symmetry meaning continuity of
the anchorage and enabling full utilisation of the fabric-to-concrete bonded area without
the adverse effects of development length of the bidirectional fibers. This was
accomplished by the construction of steel angle slotted movement joints, the details of
which are presented in figure 6.1 and 6.2. Each angle (100x100x10mm) contained 2 no.
°x 11mm slots which were placed between two greased steel plates with 10mm holding
dowels to create the movement joint. The result of this symmetric boundary was that the
effects of the patch anchors used to anchor multiple laminates spaced at 420, 320 and
220mm apart could be investigated by simulating symmetry.
In addition, 2 different patch anchor lengths: 300 mm in types (1, 3 and 4) and 250mm
(type 2) were investigated in an effort to determine the minimum anchorage length
required. With the above criteria in mind, a control specimen together with 4 types of
anchorage specimens were designed, the properties of which are presented in table 6.1
and figure 6.1.
Chapter 6 – Experimental Investigation into the Size Effect of Bidirectional Fiber Patch
Anchors
113
Type Ref Anchor length, mm
Anchor width, mm
0 0.1 control control 0.2 control control 0.3 control control
1 1.1 300 420 1.2 300 420
2 2.1 250 420 2.2 250 420
3 3.1 300 320 3.2 300 320 3.3 300 320 3.4 300 320
4 4.1 300 220 4.2 300 220 4.3 300 220 4.4 300 220
Table 6.1 – Summary of test specimens constructed in experimental program
Chapter 6 – Experimental Investigation into the Size Effect of Bidirectional Fiber Patch
Anchors
114
Figure 6.1 – Stage 2, specimen summary
Chapter 6 – Experimental Investigation into the Size Effect of Bidirectional Fiber Patch
Anchors
115
Figure 6.2 – Slotted movement joints component summary
Chapter 6 – Experimental Investigation into the Size Effect of Bidirectional Fiber Patch
Anchors
116
6.2.2 Specimen preparation
The specimens were prepared using the same techniques adopted in stage 1 to ensure
consistency in the experimental results. The surface of the concrete blocks was
sandblasted to expose the aggregate and achieve a surface roughness of approximately
1.5 mm. The major steps in the application process are summarised in figure 6.3 and
commenced with the application of a primer. Once the primer reached a tacky state,
application of the first layer of bidirectional fabric commenced. The fabric was
thoroughly impregnated with saturant and any voids within the bond line were removed
with the assistance of a hard rubber roller. The FRP laminate was applied to the surface
of the first layer of bidirectional fabric, together with 50mm adhesive tapers depicted in
figure 6.3 (b) to achieve a smooth transition of the final layer of bidirectional fabric
sheet. Finally, the second layer of bidirectional fiber was placed and 7 days of curing at
a temperature of above 25 degrees Celsius.
(a)
(b) (c) Figure 6.3 – Summary of major stages of construction for stage 3 specimens; (a) Application of first layer of bi-direction fabric; (b) Installation of FRP laminate and creation of adhesive tapers; (c) application of final layer of bidirectional fabric and sanding prior to application of strain gauges.
Chapter 6 – Experimental Investigation into the Size Effect of Bidirectional Fiber Patch
Anchors
117
6.2.3 Experimental Setup
The near end supported (NES) single pull test configuration was adopted for direct
shear testing of each anchorage specimen. The same test rig which was used in stage 1
was also used in stage 2 of the study with some slight modifications, including a
200mm high steel chair welded to the bottom of the rig, to account for the smaller
concrete block sizes. This ensured a snug fit of the concrete blocks within the testing
rig. The rig was fastened to an MTS 1MN universal testing machine using M24 high
tensile bolts, which clamped the specimen into place. The final testing configuration is
presented in figure 6.4.
(a) (b) (c)
Figure 6.4 – Specimen testing rig details (a) configuration of test rig (front view); (b) configuration of test rig (side view); (c) Photo of specimen inside testing rig
6.2.4 Test Preparation and Material properties
Concrete blocks were reinforced nominally with 4 no.12mm diameter bars at 100mm
centres each face. The reinforcement cover used was 30mm. All specimens consisted of
a single laminate strip bonded to the surface of the concrete block with a bond length of
370mm. Table 6.2 and 6.3 summarises the material properties used as per manufacturers
specifications.
Chapter 6 – Experimental Investigation into the Size Effect of Bidirectional Fiber Patch
Anchors
118
Properties Laminate
AdhesiveSaturant Primer Units
Resin Type Epoxy Epoxy Epoxy - Specific Gravity 1.8 1.12 1.08 - Glass Transition >65 - - °C
Modulus of Elasticity 10 >3.0 0.7 GPa Lap Shear Strength to >17 - - MPa
Bond (to Concrete) >3.5 >3.5 >3.5 MPa Tensile Strength 32 >50 >12 MPa
Compressive Strength >60 >80 - MPa Flexural Strength >35 >120 >24 MPa
Full cure at: 25°C 7 7 0.208 Days 40°C 3 - 0.125 Days
Table 6.2 – Adhesives, Saturant and Primer data
Properties Bidirectional
FRP (±45°)
FRP
Laminate Units
Tensile Strength 3.79 3.3 GPaTensile Modulus 230 210 GPa Ult. Elongation 2.1 1.4 %
Density 1.8 1.56 g/cm³ Thickness 0.55 1.4 mm
Width 100 mm
Table 6.3 –FRP Properties data
6.2.5 Instrumentation and loading procedure
A series of 7 strain gauges (G1-G7) were applied to the length of the FRP laminate at 50
mm intervals. An additional 4 gauges were placed either side of the laminate (2 each
side) to measure strains in the bidirectional fibers (G8-G11). Gauges G1 and G12 were
installed at the front and back of the laminate to monitor any bending in the FRP plate
during testing indicating the presence of tilting. The strain gauge locations can be seen
in figure 6.1.
6.2.6 Image correlation photogrammetry
Optical measuring techniques are increasingly being used to provide full field
monitoring of strain and deformation over a predefined area. The technique is
Chapter 6 – Experimental Investigation into the Size Effect of Bidirectional Fiber Patch
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particularly suited to capture hot spots and stress concentrations, which typically occur
in non-homogenous and anisotropic materials. 3D image correlation photogrammetry
used a pair of high resolution digital CCD cameras in combination with a randomised
high contrast speckle pattern applied to the surface of the test specimen for the 3D
deformation measurements. Stage 2 of the experiments used the image correlation
system Vic3D (Correlated-Solutions 2010) as opposed to ARAMIS which was used in
stage 1. Specialised software processing involved first defining a subset size, which was
essentially a grid covering the entire image area. Each subset was approximately 20
pixels in size and the speckles within each subset were used to define its centroid for
monitoring and correlation with surrounding subsets. Image correlation principles were
used in the Vic3D software to precisely calculate the strains and deformations to a level
of resolution dependant largely on the speckle pattern, subset size, image contrast and
image flatness.
The overall strain accuracy that could be achieved was found to be highly sensitive very
to the speckle pattern. The pixel size, randomness, contrast, highly influence the noise
in the data, which could be observed as random, sharp increases or decreases in strain
output during loading. Various methods were trailed with the aim of producing an
optimal speck pattern. The prepared surface was then spray-painted flat white in
preparation for speck application.
The method of speckle application used in stage 1 of the experimental program utilised
black spray paint, which resulted in the speckle pattern depicted in figure 6.5 (a). Using
a half pressed nozzle, a spluttering effect of black paint was created resulting in a
randomised speckle pattern applied to the surface. Although this method provided
acceptable results, the resulting speckle pattern was less than ideal due to a fine mist of
black spray which coated the surface and reduced contrast. The sizes of the speckles
were also difficult to control, resulting in speckles smaller than the pixel size being
undetected – in addition to the presence of larger speckles which exceeded the subset
size. Other methods of speckle pattern construction were trailed including: flicking
black paint from a toothbrush and the construction of a plastic template with predrilled
holes which was applied to the surface and sprayed over with black paint. The later
method proved unsuccessful due to inadequate surface flatness causing smudges in the
Chapter 6 – Experimental Investigation into the Size Effect of Bidirectional Fiber Patch
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paint. Finally, a fine point permanent marker was used to construct the speck pattern by
hand, which is depicted in figure 6.5 (b) and applied to all specimens tested in stage 2.
The strain accuracy using this method was approximately 50 microstrain as opposed to
more than double that achieved in stage 1. The increase in preparation time was justified
by the superior accuracy and noise reduction obtained by using this technique.
(a) (b)
Figure 6.5 – Speckle pattern summary; (a) speckle pattern used in stage 1; (b) improved speckle pattern used in stage 2
Rough irregular surfaces create shadows, resulting in bias and noise in the images. As a
result, the surface of the test specimen was initially sanded to remove local indentations
and surface roughness resulting in a flat finish. Good even lighting was available in the
laboratory due to the ample natural lighting available and no external light source was
needed.
Since the specimens were loaded under displacement control at a load rate of
1mm/minute, a measuring step of 1 second was used between recording intervals. Strain
and load data was obtained from surface mounted strain gauges and image correlation
photogrammetry which could be verified with each other. The overall test set-up for the
image correlation system is depicted in figure 6.6.
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(a) (b)
(c) (d)
Figure 6.6 – Photogrammetry test set-up summary; (a) speckle patter prior to testing; (b) CCD cameras mounted; (c) CCD cameras positioned approximately 3m away from test; (d) typical strain data contour over entire specimen area.
6.3 Experimental Results
6.3.1 Quality Control Tests
6.3.1.1 Concrete compressive strength The concrete material was supplied, pre-mixed from a local supplier. Compressive tests
were carried out in accordance with AS 1012.9 (1999). A total of 12 concrete cylinders
were tested to assess the concrete compressive strength. After 53 days curing at room
temperature, the average compressive strength of the concrete was 69.2MPa. The
cylinders were crushed approximately 1 day prior to testing and the results are
summarised in table 6.4.
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Table 6.4 – Concrete Mechanical Properties - Compressive cylinder results summary;
6.3.1.2 Concrete pull-off testing
Pull off tests are a standard test used to determine the efficiency of the bond between the
FRP and the concrete. This is accomplished by determining the tensile strength of the
concrete to which the FRP is to be bonded and assessing whether it is greater than the
minimum value, which is typically defined as 1.5 MPa according to ACI440.2-08. The
pull-off testing procedure was conducted in accordance with I.S. EN 1542 (1999) using
a sacrificial area of concrete on two separate concrete prisms. The following pull-off
testing procedure was used:
Surface preparation: The concrete surface was prepared using the same procedure used
as for FRP bonding. As a result the concrete surfaces were sandblasted, water jetted and
allowed to dry for a minimum of 7 days.
Core drilling: A core drill was fasted to the specimen so that lateral movement was not
permitted during drilling. A diamond core bit was used to drill a 50mm (internal
diameter) cylinders with an axis of 90 degrees to the surface. The drilling was carried
out to a depth of 50mm. This procedure ensured complete isolation of the area of
No. Concrete Cylinders
MPa1 69.52 2 68.75 3 68.12 4 67.48 5 69.39 6 69.14 7 70.92 8 68.12 9 71.30 10 70.66 11 68.12 12 69.39
Average stress, MPa 69.24
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concrete bonded to the dolly from the outer area of concrete, such that all of the stresses
were confined to the 50mm diameter cylinder of concrete.
Applying the dolly: The dollies were cleaned using abrasive paper and later degreased
using acetone prior to application. The adhesive was prepared according the
manufacturers guidelines and a thin layer was applied to the surface of the specimen so
that the adhesive formed a uniform layer between the dolly and the substrate. The
50mm aluminium dollies were placed on the core face so that the centre of the dolly
coincided with the centre of the core. Light pressure was applied to the dolly in order to
expel air while simultaneously removing and excess saturant. The adhesive was allowed
to cure for 7 days prior to testing.
Applying the load: The load was applied using a Proceq Dyna pull-off tester shown in
figure 4. The load was applied continuously at an even rate of 0.05 MPa/s until failure
occurred. Figure 6.7 depicts the pull-off testing in progress.
Test results: A total of 8 pull-off tests were conducted and the results are shown in table
6.5, which indicated an average tensile strength of 5.02 MPa. Failure was expected to
occur along the weakest plane in the system, which could be either through the
adhesive, concrete, the interface between the dolly and the adhesive or the interface
between adhesive and the concrete. The results indicated that in all cases, failure
occurred within the concrete a few millimetres below the concrete surface.
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(a) (b)
(c) (d)
(c) (b) Figure 6.7 – Summary of pull-off testing in progress and upon completion; (a) aluminium dolly applied prior to testing; (b) pull-off test depicting failure within concrete; (c) pull-off test in progress; (d) pull-off test completed.
No. Pulloff test results
MPa 1 5.54 2 4.75 3 4.57 4 NA 5 4.95 6 4.85 7 5.54 8 4.93
Average 5.02
Table 6.5 – Concrete Mechanical Properties - Pull-off test results summary
Chapter 6 – Experimental Investigation into the Size Effect of Bidirectional Fiber Patch
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6.3.1.3 Adhesive properties
Bulk adhesive moulds were filled with mixed laminate adhesive and saturant to make
tensile dumbbell and specimens. All specimens were left to cure for 7 days. All tensile
bulk adhesive specimens were tested to ISO 18280:2010. A universal testing machine,
applied a tensile loading at a load rate of 0.5 mm/min. A total of three samples were
tested and the results averaged a tensile strength of 24.8 MPa. The results are shown in
table;;6.6.
MBRACE Saturant
Tensile Strength Ultimate strain
Elastic Modulus
Poissons ratio
MPa MPa MBS1 25.8 0.0143 2025 0.3 MBS2 25.1 0.0159 2025 0.34 MBS3 23.6 0.0137 2025 0.33 Mean 24.8 0.0146 2025 0.32
Table 6.6 – Adhesive Mechanical Properties - Tensile dumbbell results
6.3.1.4 FRP Laminate properties
Based on a procedure similar to that used in stage 1 of the experimental program, the
tensile strength and elastic modulus of the FRP laminates were verified using three
laminate coupon tests in order to verify the manufacturers quoted material properties.
The FRP elastic modulus was determined using testing procedures in accordance with
ASTM: D 3039 (2000). Each test coupon had an overall length of 200 mm and average
width of 50 mm. A single strain gauge was installed at the centre of the specimen and
the strain reading was used to find the modulus of the FRP. The results indicated a mean
elastic modulus of approximately 210 GPa which verified the manufacturers value.
6.3.2 Failure Modes
Both control specimens failed by debond within the concrete cover zone within the
initial 50mm of bond length. Further along the laminate the failure plane shifted to the
interface between the concrete and adhesive, refer figure 6.8. Two alternative failure
Chapter 6 – Experimental Investigation into the Size Effect of Bidirectional Fiber Patch
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126
modes were observed in the specimens anchored with bidirectional fibers. All anchored
specimens exhibited partial debonding between the concrete and adhesive, over the
initial 50mm unanchored bond length at a load level of 90-100 kN. Load was sustained
as stresses were dispersed further along the laminate and through the bidirectional
fibers. The final failure modes observed were: (1) complete debonding of the
sandwiched laminate and bidirectional fabric structure from the concrete block, refer
figure 6.9; or (2) slippage of the laminate from between the two layers of bidirectional
fibers, refer figure 6.10.
(a) (b) (c)
Figure 6.8 – Control Specimen failure summary; (a) Concrete-adhesive separation failure (left view); (b) Back of laminate showing a combination of advesive concrete separation failure and concrete wedge failure; (c) Concrete-adhesive separation failure (right view)
Chapter 6 – Experimental Investigation into the Size Effect of Bidirectional Fiber Patch
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(b) (b)
Figure 6.9 – Patch Anchor debond (Mode I); (a) front view; (b) patch anchor pull-off failure depicting failure between saturant and the concrete
(c) (b)
Figure 6.10 – Patch Anchor debond (Mode II); (a) laminate slippage; (b) close up view It was observed that specimens with a higher concentration of aggregate at the bond interface fail by laminate slippage, whereas specimens with a lower concentration of aggregate failed by complete patch anchor debonding. 6.3.3 Overview
Table 6.7 summarises the failure loads and maximum FRP elongations reached in all
specimens tested. In the following tables and figures, reference is made to V3D
(Photogrammetry) and SG (strain gauge). These refer to the two data acquisition
techniques used in the experimental program.
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Specimen
Width of
Patch
Anchor
(mm)
Length of
Patch
Anchor
(mm)
Exp
Failure
Load
(kN)
Exp Max
Laminate
strain
( ) 0.1 NA NA 83.8 2875 0.2 NA NA 81.0 3062 0.3 NA NA 83.0 3100 1.1 400 300 131.0 4406 1.2 400 300 140.2 4922 2.1 400 250 111.0 3819 2.2 400 250 128.1 4328 3.1 300 300 151.6 5378 3.2 300 300 138.5 4801 3.3 300 300 158.8 5600 3.4 300 300 139.1 5091 4.1 200 300 140.6 4950 4.2 200 300 119.9 4504 4.3 200 300 112.5 4124 4.4 200 300 123.6 4514
Table 6.7 – Results summary
It is apparent that control specimens 0.1-0.3 exhibited fairly consistent failure loads and
elongations prior to debond with only a 3.4% variation between specimens (81-83.8
kN). All specimens which included a form of patch anchorage exhibited improvements
in strength and deformation compared to their unanchored counterparts. Since
specimens 2.1 and 2.2 were designed with a lower patch anchorage length (250mm),
they tended to exhibit laminate slippage at a lower load rather than patch anchor
debond. Laminate slippage was an unexpected failure mode and occurred due to the
reduced contact area between the laminate and each layer of bidirectional fabric. Other
contributing factors which causes laminate slippage at a lower load was the reduced
laminate width adopted in stage 2 of the study (100mm) which further contributed to a
loss of available bond area between the FRP laminate and the bidirectional fibers.
Specimens that adopted an anchorage length of 300mm (types 1, 3&4) failed either by
laminate slippage at a higher load or patch anchor debond. Due to the tendency for
premature slippage to occur when using lower anchorage lengths it was decided that
Chapter 6 – Experimental Investigation into the Size Effect of Bidirectional Fiber Patch
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129
further investigations into 250mm long anchorages be halted in favour of 300mm for
use in future study.
Specimens which used 300mm anchor lengths and failed by laminate slippage reached
load levels prior to slippage between 138.5-151.6 kN (9.5% variation). Debonding of
the entire fabric-laminate structure from the concrete (herein patch anchor debond) was
observed at similarly high levels, 131.0-140.2 kN for anchor widths of 320mm and
greater and at a lower load of 119.9 kN for 220mm wide anchors.
In general, type 4 specimens which used a 220mm anchor width were found to exhibit
lower failure loads than specimens which used an anchor width of 320mm or higher.
The reduced anchor width also resulted in a shifting of the predominating failure mode
towards patch anchor debond, which further reinforces the notion of a reduced anchor
width and subsequent area, detrimentally affecting the strength of the joint for anchor
widths of 220mm or lower.
It is useful to compare the resulting laminate strains reached in the experimental
program with permissible strains in design guidelines such as the (ACI 440.2R-08 2008)
for shear strengthened members. ACI440.2 section 11.4.1.2 follows the procedure of
applying a bond reduction coefficient to the FRP rupture strain to account for debonding
of the FRP from the concrete before loss of aggregate interlock of the section. Using
this procedure for the control specimen, the effective strain the in FRP was calculated to
be 2365 , based on the laminate properties shown in table 6.3. The predicted strain
value was 21.5% lower than the average of that obtained for the control specimens,
which is reasonable considering the design guideline is providing a lower bound
prediction.
Interestingly, the effective design strain as predicted by ACI440.2 was 16.9% of fiber
rupture strain, further reinforcing the low level of material utilisations currently being
implemented in practice and the potential for improvement through the use of anchorage
systems. Since all patch anchors, which used an anchorage length of 300mm, exceeded
4406-5378 prior to failure, material utilization could potentially be increased up to
31-38% when patch anchors are provided.
Chapter 6 – Experimental Investigation into the Size Effect of Bidirectional Fiber Patch
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130
6.3.4 Load Deformation curves
Excellent correlations are observed between the two data acquisition techniques. Less
noise can be observed in the photogrammetry measurements compared with those
obtained in stage 1 of the experimental program. This was attributed to the refined
speckle pattern used in stage 2, which provided an improvement in strain accuracy in
addition to a reduction in noise.
The load vs strain data for the control specimens is presented in figure 6.11. With the
exception of strain gauge G2 for specimen 0.2 (which malfunctioned prior testing), all
gauges performed well until failure. The stages of delamination can be observed from
the load deformation curves by examining the strains at each monitoring point location.
Debond can be observed by a reduction in gradient of the load-strain curve caused by a
loss of interfacial stiffness due to transverse micro-cracking. Sudden debond can be
observed by an abrupt increase in strain at a sustained load which can be observed in
figures 6.12 (b) at loads of 70 kN.
The load vs strain graphs in figure 6.12 for specimen type 1, showed a linear strain
increase during loading for gauges G1 and G2. This confirmed uniform loading of the
specimen. Subtle strain increases occurred across gauges G4 and G5 at approximately
70 kN indicating the onset of debonding. No appreciable strains were recorded by
gauges G5 and G6 until a load of approximately 120 kN was reached. All strain gauges
performed well during loading, however a slippage of 1000 was observed in gauge
G2 (for specimen 1.2) at approximately 100 kN.
The strain patterns with respect to loading are presented in figure 6.13 for specimen
type 2. Some minor slippage was observed in strain gauge G3 (Spec 2.2) at 110 KN,
however all other strain gauges performed well. It was also observed that strain gauges
G6 and G7 (further away from the loaded edge) exhibited a lower level of strain prior to
failure, when compared to anchor type 1 at the same location. The lower strain readings
were attributed to the reduced anchorage length and laminate slippage predominating
the failure for both type 2 specimens. The same observation can be made in other
anchorage type 3 and 4 where laminate slippage predominated, the maximum strains
Chapter 6 – Experimental Investigation into the Size Effect of Bidirectional Fiber Patch
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131
reached in gauges G6 and G7 were minimal. Figure 6.14 presents the strain results for
anchorage type 3, were data was lost for strain gauge G6 for specimen types 3.1-3.4, in
addition to gauge G4 for specimen type 3.2. The loss of data was not problematic, since
good correlations are present with the photogrammetry data. Strain gauge G1 for
specimen type 3.1 also malfunctioned as can be seen by the excessive strain readings of
over 8000 . The data for this gauge is shown indicatively, however the
photogrammetry measurements are expected to provide a more reliable strain reading at
this location.
The strain-load data for anchorage type 4 is presented in figure 6.15. Some strain gauge
temporary slippage was observed in specimens 4.1, 4.2 and 4.4, which is apparent by
the sudden reduction in strain at a sustained level of load. At such locations the
photogrammetry measurements are looked to for the correct strain readings.
(a) (b)
(c)
Figure 6.11 - Load vs strain distribution; (a) Spec 0.1; (b) Spec 0.2; (c) Spec 0.3
0
10
20
30
40
50
60
70
80
90
0 500 1000 1500 2000 2500 3000
Load
(kN)
Micro strain ( )
G1 (SG)
G2 (SG)
G3 (SG)
G4 (SG)
G5 (SG)
G6 (SG)
G7 (SG)
G1 (V3D)
G2 (V3D)
G3 (V3D)
G4 (V3D)
G5 (V3D)
G6 (V3D)
G7 (V3D)
0
10
20
30
40
50
60
70
80
90
0 500 1000 1500 2000 2500 3000
Load
(kN)
Micro strain ( )
G1 (SG)
G3 (SG)
G4 (SG)
G5 (SG)
G6 (SG)
G7 (SG)
G1 (V3D)
G2 (V3D)
G3 (V3D)
G4 (V3D)
G5 (V3D)
G6 (V3D)
G7 (V3D)
0
10
20
30
40
50
60
70
80
90
0 500 1000 1500 2000 2500 3000
Load
(kN)
Micro strain ( )
G1 (SG)G2 (SG)G3 (SG)G4 (SG)G5 (SG)G6 (SG)G7 (SG)G1 (V3D)G2 (V3D)G3 (V3D)G4 (V3D)G5 (V3D)G6 (V3D)G7 (V3D)
Chapter 6 – Experimental Investigation into the Size Effect of Bidirectional Fiber Patch
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132
(a) (b)
Figure 6.12 - Load vs strain distribution; (a) Spec 1.1; (b) Spec 1.2
(a) (b)
Figure 6.13 - Load vs strain distribution; (a) Spec 2.1; (b) Spec 2.2
0
20
40
60
80
100
120
140
0 1000 2000 3000 4000 5000
Load
(kN)
Micro strain ( )
G1 (SG)G2 (SG)G3 (SG)
G4 (SG)G5 (SG)
G6 (SG)G7 (SG)
G1 (V3D)G2 (V3D)G3 (V3D)
G4 (V3D)G5 (V3D)
G6 (V3D)G7 (V3D)
0
20
40
60
80
100
120
140
0 1000 2000 3000 4000 5000
Load
(kN)
Micro strain ( )
G1 (SG)
G2 (SG)
G3 (SG)
G4 (SG)
G5 (SG)
G6 (SG)
G7 (SG)
G1 (V3D)
G2 (V3D)
G3 (V3D)
G4 (V3D)
G5 (V3D)
G6 (V3D)
G7 (V3D)
0102030405060708090
100110120
0 1000 2000 3000 4000
Load
(kN)
Micro strain ( )
G1 (SG)
G2 (SG)
G3 (SG)
G4 (SG)
G5 (SG)
G6 (SG)
G7 (SG)
G1 (V3D)
G2 (V3D)
G3 (V3D)
G4 (V3D)
G5 (V3D)
G6 (V3D)
G7 (V3D)0
20
40
60
80
100
120
140
0 1000 2000 3000 4000 5000
Load
(kN)
Micro strain ( )
G1 (SG)
G2 (SG)
G3 (SG)
G4 (SG)
G5 (SG)
G6 (SG)
G7 (SG)
G1 (V3D)
G2 (V3D)
G3 (V3D)
G4 (V3D)
G5 (V3D)
G6 (V3D)
G7 (V3D)
Chapter 6 – Experimental Investigation into the Size Effect of Bidirectional Fiber Patch
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133
(a) (b)
(c) (d)
Figure 6.14 - Load vs strain distribution; (a) Spec 3.1; (b) Spec 3.2; (c) Spec 3.3; (d) Spec 3.4
(a) (b)
0
20
40
60
80
100
120
140
160
0 2000 4000 6000 8000
Load
(kN)
Micro strain ( )
G1 (SG)G2 (SG)G3 (SG)G4 (SG)G5 (SG)G7 (SG)G1 (V3D)G2 (V3D)G3 (V3D)G4 (V3D)G5 (V3D)G6 (V3D)G7 (V3D)
0
20
40
60
80
100
120
140
160
0 2000 4000 6000
Load
(kN)
Micro strain ( )
G1 (SG)G2 (SG)G3 (SG)G4 (SG)G5 (SG)G7 (SG)G1 (V3D)G2 (V3D)G3 (V3D)G4 (V3D)G5 (V3D)G6 (V3D)G7 (V3D)
0
20
40
60
80
100
120
140
160
0 1000 2000 3000 4000 5000 6000
Load
(kN)
Micro strain ( )
G1 (SG)G2 (SG)G3 (SG)G4 (SG)G5 (SG)G7 (SG)G1 (V3D)G2 (V3D)G3 (V3D)G4 (V3D)G5 (V3D)G6 (V3D)G7 (V3D) 0
20
40
60
80
100
120
140
0 1000 2000 3000 4000 5000 6000
Load
(kN)
Micro strain ( )
G1 (SG)G2 (SG)G3 (SG)G4 (SG)G5 (SG)G7 (SG)G1 (V3D)G2 (V3D)G3 (V3D)G4 (V3D)G5 (V3D)G6 (V3D)G7 (V3D)
0
20
40
60
80
100
120
140
0 1000 2000 3000 4000 5000 6000
Load
(kN)
Micro strain ( )
G1 (SG)
G2 (SG)
G3 (SG)
G4 (SG)
G5 (SG)
G6 (SG)
G7 (SG)
G1 (V3D)
G2 (V3D)
G3 (V3D)
G4 (V3D)
G5 (V3D)
G6 (V3D)
G7 (V3D) 0102030405060708090
100110120130
0 1000 2000 3000 4000 5000
Load
(kN)
Micro strain ( )
G1 (SG)
G2 (SG)
G3 (SG)
G4 (SG)
G5 (SG)
G6 (SG)
G7 (SG)
Chapter 6 – Experimental Investigation into the Size Effect of Bidirectional Fiber Patch
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134
(c) (d)
Figure 6.15 - Load vs strain distribution; (a) Spec 4.1, (b) Spec 4.2, (a) Spec 4.3, (b) Spec 4.4
6.3.5 FRP strain distributions along length of laminate
The strain distributions along the length of the laminates at different stages of loading
are presented in figures 6.16 to 6.20 below. The horizontal axis is measured from strain
gauge G1, which means the edge of the concrete block corresponds to a distance of
150mm along the laminate.
The photogrammetry measurements provided a continuous strain profile along the
laminate length and correlated well with the strain gauge results. In many of the figures
presented, a dip in strain is observed at approximately 150mm along the length of the
laminate where the laminate first makes contact with the concrete block. Such
behaviours could not be observed in the strain gauge measurements, since no strain
gauge was installed at this location. The dip in strain could be attributed to the
resistance provided by concrete bond in the transition between the bonded and
unbonded portions of the laminate.
At lower levels of loading, prior to the occurrence of debonding, the strain distribution
along the length of the laminate is observed to follow a non-linear descending trend.
This observation is consistent with current bond strength prediction models (Maeda et
al. 1997; Chen and Teng 2001);. At the commencement of debonding the strain is seen
0102030405060708090
100110120
0 1000 2000 3000 4000 5000
Load
(kN)
Micro strain ( )
G1 (SG)G2 (SG)G3 (SG)G4 (SG)G5 (SG)G7 (SG)G1 (V3D)G2 (V3D)G3 (V3D)G4 (V3D)G5 (V3D)G6 (V3D)G7 (V3D) 0
102030405060708090
100110120130
0 1000 2000 3000 4000 5000
Load
(kN)
Micro strain ( )
G1 (SG)G2 (SG)G3 (SG)G4 (SG)G5 (SG)G7 (SG)G1 (V3D)G2 (V3D)G3 (V3D)G4 (V3D)G5 (V3D)G6 (V3D)G7 (V3D)
Chapter 6 – Experimental Investigation into the Size Effect of Bidirectional Fiber Patch
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135
to straighten between the debonded portions of the laminate. As a result, debonding can
be observed as a gradual flattening of the strain along the laminate with increasing
loading. This further indicates the stress transfer along the laminate strip as load is
applied and debonding propagates.
Local increases or reductions in strain, causing a reversal in curvature gradient, can be
observed along the length of the laminate which deviate locally from the nonlinear
descending trend. This observation is most apparent in specimens 1.2 and 1.2, prior to
failure, where troughs and valleys are apparent in the strain distribution. A likely
explanation is the non-uniform bond strength/stiffness along the bond line creating
regions of higher stress and local cracking and spalling of the concrete, in the vicinity of
the FRP bond affecting the flatness of the laminate.
(b) (b)
(c)
Figure 6.16 - Strain vs distance along Laminate; (a) Spec 0.1; (b) Spec 0.2; (c) Spec 0.3
0
500
1000
1500
2000
2500
3000
3500
75 150 225 300 375
Micro
strain
()
Distance From Gauge G1 (mm)
30kN(V3D)
50kN(V3D)
70kN(V3D)
84kN(V3D)
30kN (SG)
50kN (SG)
70kN (SG)
84kN (SG)
0
500
1000
1500
2000
2500
3000
3500
75 150 225 300 375
Micro
strain
()
Distance From Gauge G1 (mm)
30kN(V3D)
50kN(V3D)
70kN(V3D)
80kN(V3D)
30kN (SG)
50kN (SG)
70kN (SG)
84kN (SG)
0
500
1000
1500
2000
2500
3000
3500
75 150 225 300 375
Micro
strain
()
Distance From Gauge G1 (mm)
30kN(V3D)50kN(V3D)70kN(V3D)83kN(V3D)30kN (SG)50kN (SG)70kN (SG)84kN (SG)
Chapter 6 – Experimental Investigation into the Size Effect of Bidirectional Fiber Patch
Anchors
136
(a) (b)
Figure 6.17 - Strain vs distance along Laminate; (a) Spec 1.1; (b) Spec 1.2;
(a) (b)
Figure 6.18 - Strain vs distance along Laminate; (a) Spec 2.1; (b) Spec 2.2;
(a) (b)
0500
1000150020002500300035004000450050005500
75 150 225 300 375
Micro
strain
()
Distance From Gauge G1 (mm)
20kN(V3D)60kN(V3D)100kN(V3D)130kN(V3D)20kN (SG)60kN (SG)100kN (SG)130kN (SG)
0500
1000150020002500300035004000450050005500
75 150 225 300 375
Micro
strain
()
Distance From Gauge G1 (mm)
20kN(V3D)60kN(V3D)100kN(V3D)140kN(V3D)20kN (SG)60kN (SG)100kN (SG)140kN (SG)
0
500
1000
1500
2000
2500
3000
3500
4000
75 150 225 300 375
Micro
strain
()
Distance From Gauge G1 (mm)
30kN(V3D)60kN(V3D)90kN(V3D)107.2kN(V3D)30kN (SG)60kN (SG)90kN (SG)107.2kN (SG)
0500
100015002000250030003500400045005000
75 150 225 300 375
Micro
strain
()
Distance From Gauge G1 (mm)
20kN(V3D)60kN(V3D)100kN(V3D)128kN(V3D)20kN (SG)60kN (SG)100kN (SG)128kN (SG)
0
1000
2000
3000
4000
5000
6000
7000
8000
75 150 225 300 375
Micro
strain
()
Distance From Gauge G1 (mm)
40kN(V3D)80kN(V3D)120kN(V3D)151kN(V3D)40kN (SG)80kN (SG)120kN (SG)151kN (SG)
0500
1000150020002500300035004000450050005500
75 150 225 300 375
Micro
strain
()
Distance From Gauge G1 (mm)
20kN(V3D)60kN(V3D)100kN(V3D)138kN(V3D)20kN (SG)60kN (SG)100kN (SG)138kN (SG)
Chapter 6 – Experimental Investigation into the Size Effect of Bidirectional Fiber Patch
Anchors
137
(c) (d)
Figure 6.19 - Strain vs distance along Laminate; (a) Spec 3.1; (b) Spec 3.2; (c) Spec 3.3; (d) Spec 3.4;
(a) (b)
(c) (d)
Figure 6.20 - Strain vs distance along Laminate; (a) Spec 4.1; (b) Spec 4.2; (c) Spec 4.3; (d) Spec 4.4.
0
1000
2000
3000
4000
5000
6000
75 150 225 300 375
Micro
strain
()
Distance From Gauge G1 (mm)
40kN(V3D)80kN(V3D)120kN(V3D)157kN(V3D)40kN (SG)80kN (SG)120kN (SG)157kN (SG)
0
1000
2000
3000
4000
5000
6000
75 150 225 300 375 450
Micro
strain
()
Distance From Gauge G1 (mm)
40kN(V3D)
80kN(V3D)
110kN(V3D)
138.8kN(V3D)
40kN (SG)
80kN (SG)
110kN (SG)
138kN (SG)
0500
1000150020002500300035004000450050005500
75 150 225 300 375 450
Micro
strain
()
Distance From Gauge G1 (mm)
20kN(V3D)60kN(V3D)100kN(V3D)140kN(V3D)20kN (SG)60kN (SG)100kN (SG)140kN (SG)
0
500
1000
1500
2000
2500
3000
3500
4000
75 150 225 300 375
Micro
strain
()
Distance From Gauge G1 (mm)
40kN (SG)60kN (SG)70kN (SG)80kN (SG)90kN (SG)100kN (SG)120kN (SG)
0500
100015002000250030003500400045005000
75 150 225 300 375 450
Micro
strain
()
Distance From Gauge G1 (mm)
40kN(V3D)70kN(V3D)90kN(V3D)112kN(V3D)40kN (SG)70kN (SG)90kN (SG)112kN (SG)
0500
100015002000250030003500400045005000
75 150 225 300 375
Micro
strain
()
Distance From Gauge G1 (mm)
20kN(V3D)
60kN(V3D)
100kN(V3D)
123kN(V3D)
20kN (SG)
60kN (SG)
100kN (SG)
123kN (SG)
Chapter 6 – Experimental Investigation into the Size Effect of Bidirectional Fiber Patch
Anchors
138
6.3.6 Strain in Bidirectional fibers
Measurements were conducted using surface mounted strain gauges and
photogrammetry to determine the strains in the bidirectional fibers either side of the
laminate at progressive stages of loading. It should be noted that the strains are
orientated at ±45 degrees from the longitudinal axis, resulting in an orientation which is
parallel to the direction of the bidirectional fibers. The results are summarized in figure
6.21 to 6.24 for each of the 4 types of specimens tested.
In general, the specimens which failed by debonding of the bidirectional fiber sheet,
showed a greater engagement of bidirectional fiber strains at a distance of 50mm away
from the laminate edge when compared to the specimens which failed by laminate
slippage. It is also clear that in all cases, the bidirectional fiber strains are concentrated
within the first 50mm length away from the laminate edge and follow a non-linear
descending trend, reducing to zero at a distance of 100-150mm away from the laminate
edge.
The gradient of the curves are indicative of the level of bond stress within the concrete.
The bond stresses are observed to reduce with reducing gradient as the strains in the
bidirectional fibers approached zero. These results provide insight into the minimum
spacing that patch anchored FRP laminates may be placed beside one another without a
reduction in anchorage pull-off strength. Since the strength of the anchorage is
dependent on the interfacial bond stresses between the patch anchor and concrete being
exceeded, it is expected that a laminate spacing less than 250mm would result in
superposition of bond stresses, due to strain interactions between adjacent laminates. To
mitigate such influences, a general recommendation is that: where continuous
anchorage of multiple laminates is adopted, laminates should not be spaces less than
250mm apart without incurring a loss in anchorage strength.
Chapter 6 – Experimental Investigation into the Size Effect of Bidirectional Fiber Patch
Anchors
139
(a)
(b) Figure 6.21 – Strain of 45º Bidirectional FRP either side of laminate; (a) Spec 1.1, (b) Spec 1.2
(a)
0
500
1000
1500
2000
2500
3000
3500
4000
200 150 100 50 0 50 100 150 200
Microstrain(
)
Distance ±45º from centre of laminate (mm)
40kN (SG)
60kN (SG)
80kN (SG)
100kN (SG)
110kN (SG)
130kN (SG)
40kN (V3D)
60kN (V3D)
80kN (V3D)
100kN (V3D)
110kN (V3D)
130kN (V3D)
0
1000
2000
3000
4000
5000
6000
200 150 100 50 0 50 100 150 200
Microstrain
()
Distance from across concrete block from centre of laminate (mm
40kN (V3D)60kN (V3D)80kN (V3D)100kN (V3D)120kN (V3D)140kN (V3D)40kN (SG)60kN (SG)80kN (SG)100kN (SG)110kN (SG)140kN (SG)
0
500
1000
1500
2000
2500
150 100 50 0 50 100 150
Microstrain
()
Distance ±45º from centre of laminate (mm)
40kN (SG)60kN (SG)80kN (SG)90kN (SG)100kN (SG)107.2kN (SG)40kN (V3D)60kN (V3D)80kN (V3D)90kN (V3D)100kN (V3D)107.2kN (V3D)
Chapter 6 – Experimental Investigation into the Size Effect of Bidirectional Fiber Patch
Anchors
140
(b)
Figure 6.22 – Strain of 45º Bidirectional FRP either side of laminate; (a) Spec 2.1, (b) Spec 2.2
(a)
(b)
0
1000
2000
3000
4000
5000
200 150 100 50 0 50 100 150 200
Microstrain
()
Distance from across concrete block from centre of laminate (mm
40kN (SG)60kN (SG)80kN (SG)100kN (SG)120kN (SG)128kN (SG)40kN (V3D)60kN (V3D)80kN (V3D)100kN (V3D)110kN (V3D)128kN (V3D)
0
500
1000
1500
2000
2500
3000
3500
4000
4500
150 100 50 0 50 100 150
Microstrain
()
Distance ±45º from centre of laminate (mm)
40kN (SG)60kN (SG)80kN (SG)100kN (SG)120kN (SG)151kN (SG)40kN (V3D)60kN (V3D)80kN (V3D)100kN (V3D)120kN (V3D)151kN (V3D)
0
1000
2000
3000
4000
200 150 100 50 0 50 100 150 200
Microstrain
()
Distance from across concrete block from centre of laminate (mm
40kN (V3D)60kN (V3D)80kN (V3D)100kN (V3D)138kN (V3D)40kN (SG)60kN (SG)80kN (SG)100kN (SG)110kN (SG)138kN (SG)
Chapter 6 – Experimental Investigation into the Size Effect of Bidirectional Fiber Patch
Anchors
141
(c)
(d)
Figure 6.23 – Strain of 45º Bidirectional FRP either side of laminate; (a) Spec 3.1, (b) Spec 3.2, (c) Spec 3.3, (d) Spec 3.4
0
1000
2000
3000
4000
5000
150 100 50 0 50 100 150
Microstrain
()
Distance from across concrete block from centre of laminate (mm
40kN (V3D)60kN (V3D)80kN (V3D)100kN (V3D)120kN (V3D)157kN (V3D)40kN (SG)60kN (SG)80kN (SG)100kN (SG)120kN (SG)157kN (SG)
0
1000
2000
3000
4000
5000
150 100 50 0 50 100 150
Microstrain
()
Distance from across concrete block from centre of laminate (mm
40kN (V3D)
60kN (V3D)
80kN (V3D)
100kN (V3D)
110kN (V3D)
138kN (V3D)
40kN (SG)
60kN (SG)
80kN (SG)
100kN (SG)
110kN (SG)
138kN (SG)
Chapter 6 – Experimental Investigation into the Size Effect of Bidirectional Fiber Patch
Anchors
142
(a)
(b)
(c)
0
500
1000
1500
2000
2500
3000
3500
4000
150 100 50 0 50 100 150
Microstrain
()
Distance ±45º from centre of laminate (mm)
40kN (SG)
60kN (SG)
80kN (SG)
100kN (SG)
120kN (SG)
140kN (SG)
40kN (V3D)
60kN (V3D)
80kN (V3D)
100kN (V3D)
120kN (V3D)
140kN (V3D)
0
500
1000
1500
2000
2500
3000
3500
4000
150 100 50 0 50 100 150
Microstrain
()
Distance from across concrete block from centre of laminate (mm
60kN (SG)70kN (SG)80kN (SG)90kN (SG)100kN (SG)120kN (SG)60kN (V3D)70kN (V3D)80kN (V3D)90kN (V3D)100kN (V3D)120kN (V3D)
0
500
1000
1500
2000
2500
3000
3500
4000
200 150 100 50 0 50 100 150 200
Microstrain
()
Distance from across concrete block from centre of laminate (mm
40kN (SG)
60kN (SG)
70kN (SG)
90kN (SG)
109kN (SG)
40kN (V3D)
60kN (V3D)
70kN (V3D)
90kN (V3D)
109kN (V3D)
Chapter 6 – Experimental Investigation into the Size Effect of Bidirectional Fiber Patch
Anchors
143
(d)
Figure 6.24 – Strain of 45º Bidirectional FRP either side of laminate; (a) Spec 4.1, (b) Spec 4.2, (c) Spec 4.3, (d) Spec 4.4
6.3.7 Experimental Bond Slip Curves
The accurate modelling of FRP-to-concrete interfaces is largely dependant on an
understanding of the interfacial bond-slip behaviour which can be obtained by the strain
measrements at two discrete locations. The shear stress of a particular location along the
length of the laminate can be found using a difference formula, while the corresponding
slip can be found by a numerical integration of the measured axial strains of the plate.
The bond slip relations are presented in equations 4.2 and 4.3.
Ref Distance along laminate from Edge of Concrete
Max Bond Stress
Corresponding slip
55mm 105mm 155mm so so so so
MPa mm MPa mm MPa mm MPa mm 0.1 SG 2.60 0.10 4.49 0.07 5.05 0.08 5.05 0.08 0.2 SG - - 7.26 0.09 2.50 0.03 7.26 0.09 0.3 SG 3.78 0.06 8.25 0.13 3.56 0.05 8.25 0.13
Average SG 3.19 0.08 6.67 0.09 3.70 0.05 6.85 0.10
Table 6.8 – Bond stress and corresponding slip results summary (type 1-4)
0
500
1000
1500
2000
2500
3000
3500
150 100 50 0 50 100 150
Microstrain
()
Distance from across concrete block from centre of laminate (mm
40kN (SG)
60kN (SG)
100kN (SG)
80kN (SG)
110kN (SG)
123kN (SG)
40kN (V3D)
60kN (V3D)
80kN (V3D)
100kN (V3D)
110kN (V3D)
123kN (V3D)
Chapter 6 – Experimental Investigation into the Size Effect of Bidirectional Fiber Patch
Anchors
144
Table 6.8 summarises the maximum bond stresses and corresponding slips obtained for
each of the three control specimens tested in the experimental program. These were
determined at three separate locations: 55, 105 and 155mm away from concrete free
edge. The maximum bond stress presented in the table was the maximum stress reached
when considering all locations along the laminate. The bond-slip curves depicted in
figure 6.25 have been generated from the strain gauge data. The results indicate a
general trend of linear ascending and non-linear descending branches of the curves. The
data provides useful insight into the strength and stiffness of the FRP-to-concrete
interface, which will be used in finite element modelling to inform the calibration of
interface models.
(a) (b)
(c)
Figure 6.25 – Bond-slip curve for interface derived from experimental data, (a) Spec 0.1; (a) Spec 0.2; (a) Spec 0.3;
0
1
2
3
4
5
6
0 0.05 0.1 0.15 0.2 0.25 0.3
Bond
Stress
(MPa
)
Slip (mm)
55
105
155
205
255 012345678
0 0.05 0.1 0.15 0.2 0.25 0.3
Bond
Stress
(MPa
)
Slip (mm)
105
155
205
0123456789
0 0.05 0.1 0.15 0.2 0.25 0.3
Bond
Stress
(MPa
)
Slip (mm)
55
105
155
205
255
Chapter 6 – Experimental Investigation into the Size Effect of Bidirectional Fiber Patch
Anchors
145
The maximum bond stresses and corresponding slips were also determined for all 4
anchorage specimens and the results are summarised in table 6.9. Although it is
unrealistic to expect that the results will be indicative of the true bond stresses in the
concrete due to the strain gauges being bonded to the outermost layer of bidirectional
fabric sheet and the likelihood of bond imperfections and slippages occuring between the
fabric, laminate and concrete. However, the values provide a reasonable estimation of the
bond stresses between the outermost layer of bidirectional fabric and the FRP laminate.
Such information is was required in the numerical simulations to define the interfacial
bond properties between the laminate and fabric.
Type Ref Distance along laminate from Edge of Concrete
Max Bond Stress
Corresp-onding
slip 55mm 105mm 155mm
so so so so MPa mm MPa mm MPa mm MPa mm
1 1.1 5.52 0.62 6.97 0.18 10.12 0.25 10.12 0.25 1.2 2.64 0.08 7.35 0.11 9.60 0.19 9.60 0.19
2 2.1 4.56 0.10 5.63 0.09 7.73 0.10 7.73 0.10 2.2 7.68 0.47 7.80 0.15 10.41 0.35 10.41 0.35
3
3.1 - 0.68 7.18 0.38 7.49 0.17 7.49 0.17 3.2 11.34 0.63 6.12 0.28 5.00 0.13 11.34 0.63 3.3 9.38 0.17 9.65 0.24 6.27 0.18 9.65 0.24 3.4 5.50 0.13 8.78 0.20 3.36 0.33 8.78 0.20
4
4.1 5.67 0.57 10.59 0.22 7.95 0.11 10.59 0.22 4.2 8.03 0.17 8.24 - 4.42 - 8.24 - 4.3 5.85 0.09 6.89 0.12 6.31 0.18 6.89 0.12 4.4 - - 11.21 0.32 6.86 0.10 11.21 0.32
Average 6.62 0.34 8.03 0.21 7.13 0.19 9.34 0.25
Table 6.9 – Bond stress and corresponding slip results summary (type 1-4)
6.4 Conclusion
Anchorage systems are a relatively new area of study with potential to improve the
performance of FRP materials bonded to concrete. This research has focused on the use
of patch anchors made of bidirectional fibers to distribute FRP-to-concrete bond stresses
within the FRP anchorage zone over a greater area of concrete. As a result, anchorage
Chapter 6 – Experimental Investigation into the Size Effect of Bidirectional Fiber Patch
Anchors
146
strength is increased, while facilitating a higher level of fiber strain prior to debond. The
results and discussions presented throughout the paper allow the following conclusions
to be made:
Patch anchorage lengths of 250mm exhibited slippage at a lower load. As a
result it was recommended that 300mm be the ideal patch anchorage length.
Specimens anchored with 300mm long anchorages (which failed due to
slippage) exhibited increases in load of 53-81%.
300mm long anchorage joints which failed by anchor debond, exhibited similar
increases in load of 56-67% for specimen type 1. However the effect of
reduction in concrete block width (200mm) used in specimen type 4, resulted in
pull-off failure at a lower load.
By examining the strain distributions within the bidirectional fibers it is
expected that laminates could be spaced as close a 250mm without any
reduction of anchorage strength.
Chapter 7 – Finite Element Modelling of Unidirectional and Bidirectional Fiber Patch
Anchors
147
7 CHAPTER 7 – FINITE ELEMENT MODELLING OF UNIDIRECTIONAL AND BIDIRECTIONAL FIBER PATCH ANCHORS
7.1 Introduction
In order to further advance the theoretical understanding of FRP patch anchors and
expand the pool of available data, it is necessary to develop and calibrate non-linear
finite element (FE) models prior to undertaking parametric studies. The data from
parametric studies can be used in the development of theoretical models and design
guidelines. Finite element simulations by (Pham 2005; Hii and Al-Mahaidi 2006; Pham
2007) have shown that the method is capable of satisfactorily modelling the pre-peak
and post-peak, non-linear, behavioural response of RC members strengthened with FRP.
The FE model developed herein is implemented in ATENA 3D (Cervenka 2007), which
is a numerical modelling package specialising in RC structures. In addition to the
definition of concrete elements capable of cracking and crushing behaviour, a non-linear
interfacial bond law was defined for the region between the FRP and concrete materials.
This law is calibrated with experimental bond-slip data. The innovative modelling and
calibration procedures adopted have resulted in a good prediction of structural response
and failure modes, the results of which can be of assistance to future researchers and can
be used in the future development of much needed design formulations for FRP
anchorage devices. The parametric studies conducted enabled extrapolation of the
experimental data over a wide range of concrete strengths.
7.2 The Proposed Finite Element Model
The FE model was implemented in ATENA 3D (Cervenka 2007) and utilised
symmetric boundary conditions through the plane of symmetry of the specimen to
reduce model size and solution time. The creation of the model included the definition
of material models for concrete, FRP laminates, FRP unidirectional and bidirectional
fabric, steel reinforcement and an interface bond law between the FRP and the concrete.
Chapter 7 – Finite Element Modelling of Unidirectional and Bidirectional Fiber Patch
Anchors
148
A summary of the FE model constructed and its respective components are depicted in
figure 7.1.
Figure 7.1 – Summary of FE model built in ATENA 3D
7.2.1 Modelling of concrete
Concrete is considered to be a quasi-brittle material capable of both cracking and
crushing behaviour under tensile and compressive stress. The compressive response of
concrete is highly non-linear, whereas in tension, the stress-strain response is described
using a softening law and the smeared crack approach incorporating the fracture energy
concept. The non-linear compressive behaviour of concrete can be captured using
numerical non-linear plasticity models. Inclusion of cracking response can be simulated
using fracture-plastic material models currently available in many FE packages.
Extensive research on the numerical modelling of concrete cracking has resulted in two
main crack models being investigated: (a) discrete crack model and (b) smeared crack
model. Discrete crack models rely on simulating discrete cracks by introducing
discontinuities within the FE mesh at element boundaries. The disadvantages of this
method are that each crack must be physically modelled and that the crack propagation
depends on the geometry and the topology of the mesh. This drawback can be overcome
Chapter 7 – Finite Element Modelling of Unidirectional and Bidirectional Fiber Patch
Anchors
149
by refinement and re-meshing but these approaches are computationally expensive
(Rabczuk et al. 2008).
In the smeared crack models, the cracked material is treated as continuous, and the
discontinuity of the displacement field caused by the crack is spread across the element
by changing the element stiffness, also known as strain softening. In general, the
smeared crack approach has grown more popular and demonstrated greater advantages
than the discrete crack method. However, the smeared crack strategy tends to spread
crack formation over a band of elements and fails to predict localised fracture. Smeared
crack models can be divided into two main categories, fixed and rotating crack models.
Fixed crack models utilise a constant crack orientation during the entire computational
process. Whereas, in the rotating crack model, the crack direction changes with load
history, corresponding to the principal directions.
The proposed material model utilised in this study is based on the smeared crack model
and refined crack band theory. The adopted fixed crack model is based on a non-linear
plasticity fracture material model utilising fracture energy and a crack opening law.
Concrete cracking was considered as part of a three stage fracture process: Uncracked,
potential crack in progress and crack opening after complete release of stress. The
compressive failure was simulated using a biaxial stress failure criterion based on
(Kupfer 1969). A reduction of compressive strength and shear stiffness after cracking
was also considered.
The input parameters for the required by the concrete material model were: Young's
modulus (Es), compressive strength (f’c,), tensile strength (fct), poisson ratio (v) and
mode 1 specific fracture energy (GF1). The chosen parameters are summarised in table
7.1. With the exception of fracture energy, all input parameters could be determined
from the experimental measurements. Recent attempts by researchers to address the
problem of premature debonding has been to anchor the ends of the externally bonded
FRP using novel anchorage devices such as: FRP U-Jackets (Al-Amery and Al-Mahaidi
2006; Pham and Al-Mahaidi 2006), metallic anchorage devices (Duthinh and Starnes
2001; Wu and Huang 2008), FRP patch anchors. FRP Anchors or spike anchors
(Micelli et al. 2010; Smith 2010) and mechanical substrate strengthening.
Chapter 7 – Finite Element Modelling of Unidirectional and Bidirectional Fiber Patch
Anchors
150
The experimental studies in chapters 5 and 6 have shown that ±45º oriented
bidirectional fabric anchorages (herein patch anchors) can result in a gradual debonding
of the FRP strengthening with the FRP-adhesive stresses being distributed across a
greater area of the concrete. The anchored joints experienced increases in strength of
93-109 % as well as loaded end slippage of 4 to 8 times above that of the unanchored
counterparts. The ±45º oriented bidirectional fabric configuration was successfully
applied in the strengthening of the West Gate Bridge in Melbourne which represents the
worlds largest application of FRP strengthening to date.
In order to further advance the theoretical understanding of FRP patch anchors and
expand the pool of available data, it is necessary to develop and calibrate FE models
prior to undertaking parametric studies. The data from parametric studies can be used in
the development of theoretical models and design guidelines. Finite element simulations
by (Pham and Al-Mahaidi 2005, 2007; Al-Mahaidi and Hii 2007) have been capable of
satisfactorily modelling the pre-peak and post-peak, non-linear, behavioural response of
RC members strengthened with FRP. The FE model developed herein is implemented in
ATENA 3D (Cervenka 2007), which is a numerical modelling package specialising in
RC structures. In addition to the definition of concrete elements capable of cracking and
crushing behaviour, a non-linear interface bond law was defined for the region between
the FRP and concrete materials. This law is calibrated with experimental bond-slip data.
The innovative modelling and calibration procedures adopted have resulted in an
accurate prediction of structural response and failure modes, the results of which can be
of assistance to future researchers and can be used in the future development of much
needed design formulations for FRP anchorage devices. The parametric studies
conducted enabled extrapolation of the experimental data over a wide range of concrete
strengths.
Chapter 7 – Finite Element Modelling of Unidirectional and Bidirectional Fiber Patch
Anchors
151
Concrete properties Model Young's modulus, Es (MPa) 39760
Mean compressive strength, f’cm 62 Characteristic tensile strength, fct 4.72
Poisson ratio, v 0.2 Specific fracture energy, GF
I 210 Shear factor coefficient 20
Table 7.1 – Summary of input parameters used in non-linear concrete model
A fracture energy of (GFI = 210 N/m) was adopted in this study was based on (Trunk
and Wittmann 1998) and using a maximum aggregate size of 20mm. The value was
chosen based on a sensitivity study on fracture energy which will be described in
section 7.1.11. The exponential softening curve was approximated in the numerical
model using a bi-linear relationship. In order to define the relationship between normal
and shear crack stiffness, a shear factor coefficient of 20 was specified based on
experimental work by Walraven (1981). This essentially defines the ratio between
normal to shear stiffness of cracked elements and the units are dimensionless.
7.2.2 Modelling FRP Patch Anchors
Type 5 FRP Patch anchors consist of loosely woven fibers orientated in the ±45°
directions and embedded within a saturant matrix. This material was simplified in the
FE model by defining a saturant base material, comprising three dimensional brick
elements of equivalent 0.86mm sheet thickness for a single layer. An isotropic linear
elastic material model with VonMises plasticity hardening was assigned to the saturant
material. The fibers were defined using smeared reinforcement in perpendicular
orientations (within the saturant elements) representing the embedded fibers. The
orientation of the smeared reinforcement could be defined such that it was ±45º to the
direction of the laminate. A fiber fraction of 19.7% in each direction was used to
replicate the orthogonal fibers. This figure was obtained by dividing the total area of
loose fibers by the area of saturant to arrive at the correct force per unit width. The FRP
laminate was modelled as an isotropic linear elastic material with properties described
in chapter 5. Unidirectional patch anchors were modelled in a similar fashion to their
bidirectional counterparts. The unidirectional fibers were also defined using a 0.86mm
Chapter 7 – Finite Element Modelling of Unidirectional and Bidirectional Fiber Patch
Anchors
152
thick homogeneous saturant base material which included embedded fibers as smeared
reinforcement with a fiber fraction of 27.36%. Material properties for the fiber and
saturant materials were based on manufacturer’s specifications.
7.2.3 Modelling steel reinforcement
The steel reinforcement was assumed to be elastic, perfectly plastic and was defined
using a bi-linear stress-strain law. Steel bars were modelled individually as embedded
reinforcements in the concrete elements, which implied a perfect bond exists between
the steel bars and concrete due to full strain compatibility. This is a reasonable
assumption, since the majority of the tensile and shear stresses during loading were
concentrated at the adhesive-to-concrete interface and the steel reinforcement was not
expected to develop any significant stress levels.
7.2.4 Modelling FRP-to-Concrete Interface
Debonding of FRP-to-concrete joints are generally governed by the interfacial bond
strength characteristics between the concrete-adhesive and adhesive-FRP, the former
being most critical. In numerical simulations, researchers have utilised several methods
for modelling this bond-interface interaction. Three different approaches have been
adopted in the literature to simulate the behaviour of the FRP–concrete interface using a
nonlinear FE model, i.e. (a) perfect bond at the interface (Cui 2009; Mohammad
Hajsadeghi 2011), (b) the use of one-dimensional nonlinear spring elements between
the adjacent concrete and adhesive layers (Luo 2011) and (c) a layer of interface
elements between the FRP and the concrete (Brena 2003; Pham and Al-Mahaidi 2006;
Freddi and Savoia 2008; Wu et al. 2009). In the first approach, a perfect bond is
assumed between the adhesive and concrete layers. Debonding is directly modelled by
simulating the cracking and crushing of concrete elements adjacent to the adhesive layer
(Hiroyuki 1997). This approach relies heavily on an accurate constitutive model for the
concrete material and the size of the concrete elements beneath the adhesive layer which
must also be sufficiently small to capture the localised debonding failure within the
concrete cover zone. Although researchers have achieved acceptable results utilising
this technique; the authors noted that for the constitutive concrete model used in this
Chapter 7 – Finite Element Modelling of Unidirectional and Bidirectional Fiber Patch
Anchors
153
study, the method was sensitive to mesh size beneath the adhesive layer resulting in
problematic model calibration. A method for modelling the FRP-to-adhesive interface
less commonly used is by introducing non-linear one-dimensional spring elements
between the adhesive and concrete layers (Luo 2012). The function of the spring
elements is to represent the shear resistance of the adhesive-concrete interface. Each
spring element is defined with a damage-type constitutive law derived from the average
interfacial shear stress and corresponding slip values. When accurately validated, the
spring element can describe the stress-slip relationship in addition to modelling the
post-peak response of the joint.
A more commonly used approach is to simulate FRP debonding by the definition of
interface elements between the FRP and the concrete. A constitutive bond-slip model or
shear traction-separation law is typically assigned to the interface elements which can
be calibrated using experimental data or by utilizing available theoretical models based
on linear or non-linear fracture mechanics (Lu et al. 2005; Yao and Teng 2005; Wang
2006; Ferracuti et al. 2007). This bond-slip relationship consists of two stages: an
initially elastic stage in which the interfacial stress increases with the slip until it
reaches the strength of the interface, and a softening stage in which interfacial stress
decreases with the slip resulting in debonding (Wang and Zhang 2008).
The present study uses interface elements as contact between two surfaces (concrete and
FRP). The constitutive relation for a general three-dimensional case is given in terms of
tractions on interface planes, relative sliding and opening displacements. The initial
failure surface corresponds to a Mohr-Couloumb condition, where after stresses violate
the shear strength limit, the surface collapses into a multi-linear softening law calibrated
with the experimental data, refer figure 7.2 (a) and (b). The input parameters for the
interface material model are summarized in table 7.2.
Chapter 7 – Finite Element Modelling of Unidirectional and Bidirectional Fiber Patch
Anchors
154
Figure 7.2 – Calibrated shear strength interface material model and cohesion softening law; (a) Numerical definition (Cervenka 2007); (b) shear-slip curve for interface derived from experimental data. Where: = interfacial shear stress, = normal stress, = friction angel, Ktt = tangential stiffness, GF
I= mode 1 fracture energy, c = cohesion
Parameter Value adopted normal stiffness, Knn 5x105
tangential stiffness, Ktt 1.2x106
Characteristic tensile 4.72Cohesion, c (MPa) 5.0
Table 7.2 – Parameters used to define interface material model
The experimental bond slip curve was used to determine the numerical parameters of,
tangential stiffness (Ktt ), cohesion (c) and friction coefficient ( ) for the shear-
displacement function in figure 7.2 (a). The multi-linear softening component was also
derived from the same experimental data and used to define the mode I fracture energy
(GFI) negating the need of a friction coefficient as there is no longer a sudden collapse
of cohesive strength to the dry friction value. Failure was replicated in the FE model by
the definition of an interface bond law between the adhesive and concrete materials and
by assigning a perfect bond between all other subsequent layers.
0
1
2
3
4
5
6
0 0.1 0.2 0.3 0.4
Bond
Stress
(MPa
)
V
175mm (GAUGE)
175mm (FEM)
(b)
Chapter 7 – Finite Element Modelling of Unidirectional and Bidirectional Fiber Patch
Anchors
155
7.2.5 Solution strategies
The Newton-Raphson method was implemented using displacement control to obtain
the numerical predictions. The step size was defined as the designated displacement per
load step and was found to have significant effects on the bond behaviour. Parametric
studies were used to determine the optimal step size of 0.005mm to facilitate optimal
convergence and capture the full spectrum of post-peak response and ultimate failure
load. It was found that larger step sizes (greater than 0.01mm) resulted in numerical
instability, solution divergence and premature failure.
The non-linear numerical analysis used an implicit solver with an iterative scheme to
bring the internal energy balance to an acceptable level of equilibrium. The degree of
error was determined by assigning convergence criteria and tolerances to errors such as:
displacement error, residual error and energy error. The final convergence criteria
assigned for the above mentioned errors were: 0.01 (displacement error), 0.01 (residual
error) and 0.0001 (energy error). These values were found to be a good compromise
between computational time and solution accuracy.
7.2.6 Element type for the concrete prism
Prelimanary studies were conducted to compare the performance the standard 8-noded
isoparametric solid brick element with a higher order sophisticated 20-node brick
element. At a lower level of load there was little difference observed between The two
models. However at higher load levels, the 20-noded element demonstrated better
numerical stability and convergence rates. Therefore, the finite element models were
created using the higher order 20 node iso-parametric solid brick elements, which were
used to model the concrete, FRP and adhesive materials (refer figure 7.3). The element
contains 120 degrees of freedom with three displacements, ux, uy and uz, and three
rotations, x, y and z at each node. A 2 x 2 x 2 Gaussian integration scheme was
adopted. A sensitivity study was performed to determine the optimal element size and
level of mesh refinement to balance accuracy with available computational resources.
Based on 4 alternative mesh sizes, which are discussed in section 7.5.1, the size of the
concrete elements chosen were 20 x 5 x 15mm below the FRP bond line.
Chapter 7 – Finite Element Modelling of Unidirectional and Bidirectional Fiber Patch
Anchors
156
Figure 7.3 – Geometry of CCIso Brick element
Figure 7.4 summarises the FE mesh used for the three models under investigation
(specimens type 0, 2 ad 3). The concrete elements were refined in thickness nearing the
FRP bond line to a minimum of 5mm and consistency in mesh size and pattern was
ensured across all specimens.
(a) (b)
(c)
Figure 7.4 – FE mesh summary: (a) Type 0 (Control); (b) Type 2 (Unidirectional patch anchor); (c) Type 5 (Bidirectional patch anchor)
Chapter 7 – Finite Element Modelling of Unidirectional and Bidirectional Fiber Patch
Anchors
157
7.3 Boundary Conditions
The FE model utilised several boundary conditions to replicate the experimental test
configuration. Symmetric boundary conditions were used through the centrelines of the
specimens in order to reduce computation time. This was achieved by fixing translations
in the x-direction and freeing all other displacements in y and z. The dimensions of the
L-shaped constraint depicted in figure 7.5 applied to the top face of the concrete was
constructed using the steel testing rig as a template and constraint was fixed in all
directions. The back face of the concrete blocks were restrained in y to prevent rotation
and the bottom face of the concrete was restrained in x and y.
Figure 7.5 – FE model boundary conditions summary.
7.4 Numerical and Experimental Results
This section summarises the experimental results of anchorage types 0 and 5 and
compares them to predictions from the numerical model. In the following tables and
figures, reference is made to AR (Photogrammetry) and SG (strain gauge). These refer
to the two data acquisition techniques used in the experimental programme. Table 7.3
summarized the maximum loads and strains reached prior to debonding for each
respective specimen. Debonding was determined experimentally by examining the load
vs strain curves at the respective monitoring points. In particular, debonding is indicated
by the sudden loss of stiffness and increase in strain shown in figures 7.7, 7.9 and 7.12.
Chapter 7 – Finite Element Modelling of Unidirectional and Bidirectional Fiber Patch
Anchors
158
Similarly, gradual vs sudden de-bonding was determined by comparing the slope of the
load vs strain curves among types 0, 2 and 5 specimens.
Type Specimen Width of
Patch
Anchor
(mm)
Length
of
Patch
Anchor
Exp
Failure
Load
(kN)
Exp Max
Laminate
strain
( )
FEM
Failure
Load
(kN)
FEM
Accuracy
(±%)
0 WG9 N.A N.A 99.6 2535 93.3 -6.75 2 WG3 400 270 138.2 3242 131.8 -4.86
WG4 400 270 142 3142 131.8 -7.74 5 WG10 400 270 213 4900 209 -1.91
WG11 400 270 236.9 5300 209 -13.35
Table 7.3 – Summary of maximum loads and FRP strains reached prior to debonding derived from experimental data (types 0, 2 & 5)
7.4.1 Type 0 – Control specimen Results
The FE model was indicative that FRP deboding occurred via a combination of
concrete-adhesive interfacial debonding and failure through the concrete elements
closest to the bond line (concrete cover separation failure). The form of debonding
achieved through the FE simulations was in agreement with experimental observations.
Concrete cracking was observed in the experiments in the vicinity of the concrete free
edge and at the interface between the concrete and the FRP. The FE model confirmed
this via the tensile stresses of the elements being exceeded resulting in the crack patterns
shown in figure 7.6.
Figures 7.7 (b) to (e) show the load vs strain curves for type 0 specimens. As can be
seen the FE model captures the overall failure load and debonding behaviour well,
failing at 92.5 kN (7.5% lower than experimental value). The strain differences
observed between the experimental and FE results recorded at 150mm away from the
concrete free edge be explained by the minor degree of tilting of the specimen during
loading which was recorded.
Chapter 7 – Finite Element Modelling of Unidirectional and Bidirectional Fiber Patch
Anchors
159
(a) 50 kN
(b) 91.6 kN
Figure 7.6 – Failure model of Control Specimen (FEM Model) depicting exaggerated deformations
(a)
0
500
1000
1500
2000
2500
0 50 100 150 200 250 300
Micro
strain
()
Distance along Laminate from gauge G1 (mm)
30kN (SG)
50kN (SG)
70kN (SG)
80kN (SG)
90kN (SG)
30kN (FEM)
50kN (FEM)
70kN (FEM)
80kN (FEM)
90kN (FEM)
Concrete Cover separation failure
Principal Tensile Stress
(MPa)
Chapter 7 – Finite Element Modelling of Unidirectional and Bidirectional Fiber Patch
Anchors
160
(b) (c)
(d) (e)
Figure 7.7 – Load vs strain distribution: (a) Strain vs distance along laminate; (b) Gauge G3; (c) Gauge G4; (d) Gauge G5; (e) Gauge G6
7.4.2 Type 5 – Bidirectional fabric specimen Results
The experimental results indicated that failure of the bidirectional fabric specimens
occurred in the interface between the adhesive and concrete layers. The result was the
complete debonding of the laminate-fabric structure from the concrete surface. The
stages of failure, interface cracking and exaggerated deformations for type 5 specimens
are schematically presented in figure 7.8 for 5 stages of loading. From a review of the
crack patters of the elements closest to the FRP-to-conceret interface, the progress of
debonding can be monitored. Crack propagation initiated at the face of the concrete
block (closest to the point of load application) and is seen to progress horizontally
0
20
40
60
80
100
120
0 500 1000 1500 2000 2500
Load
(kN)
Microstrain
G3 (SG)
G3 (AR)
G3 (FEM)0
20
40
60
80
100
120
0 500 1000 1500 2000 2500
Load
(kN)
Microstrain
G4 (SG)
G4 (AR)
G4 (FEM)
0
20
40
60
80
100
120
0 500 1000 1500 2000 2500
Load
(kN)
Microstrain
G5 (FEM)
G5 (SG)
G5 (AR)
0
20
40
60
80
100
120
0 500 1000 1500 2000 2500
Load
(kN)
Microstrain
G6 (SG)
G6 (AR)
G6 (FEM)
Chapter 7 – Finite Element Modelling of Unidirectional and Bidirectional Fiber Patch
Anchors
161
within the elements beneath the FRP bond line towards the FRP free end. Although
debonding occurred mainly by concrete cover separation failure within the zone directly
under the FRP laminate. Concrete adhesive interfacial debonding was observed to
doninate in the region under the bidirectional fiber patch, resulting in complete
separation as observed in figure 7.8 (e). The predicted modes of debonding correlated
well with the experimental observations, where complete debonding of the laminate-
fabric structure was observed at failure.
Chapter 7 – Finite Element Modelling of Unidirectional and Bidirectional Fiber Patch
Anchors
162
(a) 50 kN (b) 100 kN
(b) 150 kN (d) 209 kN
(e) Failure
Figure 7.8 – Failure model of Anchor Type 5 (FEM Model) depicting exaggerated deformations
Concrete-Adhesive
Interfacial Debonding
Concrete Cover
separation failure
Principal Tensile Stress
(MPa)
Chapter 7 – Finite Element Modelling of Unidirectional and Bidirectional Fiber Patch
Anchors
163
(a)
(a) (c)
(d) (e)
Figure 7.9 - Load vs strain distribution – Type 5 (Bidirectional fabric); (a) Strain vs Distance along FRP laminate; (b) Gauge G3; (c) Gauge G4; (d) Gauge G5; (e) Gauge G6.
0
500
1000
1500
2000
2500
3000
3500
4000
4500
5000
0 50 100 150 200 250
Micro
strain
()
Distance along Laminate from gauge G1 (mm)
50kN (FEM)75kN (FEM)100kN (FEM)125kN (FEM)150kN (FEM)175kN (FEM)211.95kN (FEM)50kN (SG)75kN (SG)100kN (SG)125kN (SG)150kN (SG)175kN (SG)211.95kN (SG)
0
50
100
150
200
250
0 1000 2000 3000 4000 5000
Load
(kN)
Micro strain ( )
G3 (FEM)G3 (SG)G3 (AR)
0
50
100
150
200
250
0 1000 2000 3000 4000
Load
(kN)
Micro strain ( )
G4 (FEM)
G4 (SG)
G4 (AR)
0
50
100
150
200
250
0 1000 2000 3000 4000
Load
(kN)
Micro strain ( )
G5 (FEM)
G5 (SG)
G5 (AR)0
50
100
150
200
250
0 1000 2000 3000 4000
Load
(kN)
Micro strain ( )
G6 (FEM)G6 (SG)G6 (AR)
Chapter 7 – Finite Element Modelling of Unidirectional and Bidirectional Fiber Patch
Anchors
164
A review of the strains along the length of the laminates presented in figure 7.9 (a)
shows good correlations with FE predictions. However, FE tended to predict lower
strains at 250mm away from the loaded face. The stiffer load vs strain response
predicted in the first half of the load vs strain curves reported in figures 7.9 (b) and 7.9
(c) is attributed to: (1) The assumptions that form the basis of the smeared crack model;
(2) the assumption of idealised bond between the concrete and adhesive layers and (3)
approximations used to define the bidirectional FRP sheet. The smeared crack model
resulted in stiffer load response, as the model failed to predict localised fracture. This
resulted in uniform cracking distributions spread over a band of elements, which can
explain the stiffer response. In addition, the modelling approach used to define the
bidirectional fabric, failed to consider the discontinuities that individual fibers create
within the saturant matrix and the non-uniformity in saturant homogeneity. As a result,
the assumptions used in the model tended to produce a stiffer fabric material. The bond
between the adhesive and concrete layers was assumed to adhere to a pre-defined
interface bond law. In reality such laws are idealistic, as variations in bond
characteristics can be affected by the local tensile strength of concrete and incongruity
in fiber application. In addition, the assumptions of perfect bond between multiple fiber
layers contribute to increasing the stiffness of the numerical model.
The FE model also predicted adequate engagement of the bidirectional fibers as
observed in figure 7.10 by comparing FE predictions with the strain values measured by
the two data acquisition systems at monitoring point locations G19-G22. It can be
observed that even near the failure load (200 kN), the strains in the bidirectional fibers
reduced to zero at a distance of 150mm away from the centre of the laminate. These
results provide insight into the minimum spacing that patch anchored FRP laminates
may be placed beside one another without expecting a reduction of anchorage pull-off
strength. Based on these results, the recommended minimum spacing between FRP
laminates should be 250mm.
Chapter 7 – Finite Element Modelling of Unidirectional and Bidirectional Fiber Patch
Anchors
165
Figure 7.10 - Type 5 (45º Bidirectional FRP) – Strain parallel to the fibers of the patch anchor (±45º) vs distance away from centre of laminate (mm)
7.4.3 Type 2 – Unidirectional fabric specimen Results
The predicted failure modes and exaggerated deformations are presented in figure 7.11
for anchorage type 2 Localized concrete-adhesive separation failure was observed
beneath the unidirectional fabric and the edges of the FRP laminate. However, concrete
cover separation failure was the predominating failure mode within the central region of
the FRP laminate. These two forms of failure were also consistently observed in the
experimental tests.
Examining the strain distributions reported in figure 7.12, three stages of debonding are
observed in two zones along the length of the laminate. The first stage of debonding
initiates between 40-80kN. This behaviour is captured by gauges G3 and G4 which
were installed in the unanchored section of the laminate and can be observed by the
gradual reduction of stiffness and corresponding increases in strains recorded in figure
7.12 (a). Similar behaviour was observed at gauges G5 and G6 (located 150-200mm
from the loading edge of the concrete block) at a higher load level (110-120kN).
Examining figure 7.12, the FE results match very closely to the experimental data
during all stages of loading. The benefits of type 5 anchor as opposed to type 2 are
clearly observed in the maximum strains reached prior to debonding, which were 60%
higher.
0
1000
2000
3000
4000
5000
6000
150 120 90 60
Microstrain
()
Distance from across concrete block from centre of laminate (mm)
200 kN (AR)
200kN (SG L)
200 kN (FEM)
Chapter 7 – Finite Element Modelling of Unidirectional and Bidirectional Fiber Patch
Anchors
166
(a) 50 kN (b) 100 kN
(c) 131.8 kN
(d) Failure
Figure 7.11 – Failure model of Anchorage Type 3 (FEM Model) depicting exaggerated deformations
Concrete-Adhesive
Interfacial Debonding
Concrete Cover
separation failure
Principal Tensile Stress
(MPa)
Chapter 7 – Finite Element Modelling of Unidirectional and Bidirectional Fiber Patch
Anchors
167
(a)
(b) (c)
(d) (e)
Figure 7.12 - Load vs strain distribution, Type 2 (Unidirectional fabric); (a) Gauge G3; (b) Gauge G4; (c) Gauge G5; (d) Gauge G6
0
500
1000
1500
2000
2500
3000
3500
0 50 100 150 200 250 300
Micro
strain
()
Distance along Laminate from strain gauge G1(mm)
40kN (FEM)
80kN (FEM)
110kN (FEM)
130kN (FEM)
40kN (SG)
80kN (SG)
110kN (SG)
137kN (SG)
0
20
40
60
80
100
120
140
160
0 1000 2000 3000 4000
Load
(kN)
Micro strain ( )
G3 (AR)
G3 (SG)
G3 (FEM)
0
20
40
60
80
100
120
140
160
0 1000 2000 3000 4000
Load
(kN)
Micro strain ( )
G4 (FEM)
G4 (SG)
G4 (AR)
0
20
40
60
80
100
120
140
160
0 500 1000 1500 2000 2500 3000
Load
(kN)
Micro strain ( )
G5 (AR)
G5 (SG)
G5 (FEM)0
20
40
60
80
100
120
140
160
0 500 1000 1500 2000 2500 3000
Load
(kN)
Micro strain ( )
G6 (SG)
G6 (AR)
G6 (FEM)
Chapter 7 – Finite Element Modelling of Unidirectional and Bidirectional Fiber Patch
Anchors
168
7.5 Parametric studies
The FE models described above produced analogous results to the experiments. The
models have been ascertained to be valid. However, several parameters within the
material models were selected from a range of possible options so that the best fitted
results could be found. Therefore, in the sections to follow, a parametric study is carried
out to investigate the sensitivity of the numerical results to a number of parameters such
as: mesh size, concrete fracture energy and adhesive stiffness. The sensitivity study was
performed for the control specimen only as the interface behaviour and mechanisms of
load transfer between the FRP-to-concrete and were the same for all subsequent
specimens.
7.5.1 Sensitivity to mesh size
To study the model sensitivity to mesh sizes, three meshes were generated and they are
listed in Table 7.4 where (x) represents the width across the laminate, (y) the thickness
of concrete elements below the laminate and (z) the length of elements along the
laminate.
Model Mesh size, x, y, z (mm)
Mesh 1 10 x 2.5 x 15 Mesh 2 10 x 5 x 15 Mesh 3 20 x 5 x 25 Mesh 4 20 x 15 x 25
Table 7.4 - Mesh size variations
A comparison of the load vs strain predictions are presented in figure 7.13 for strain
gauge locations G3 to G5. It is apparent that the smaller the mesh size the more accurate
predictions can be obtained. The predictions of the FRP behaviour in terms of load,
strain, bond stress and slip are very close to the experimental values for mesh sizes 1-3.
Since only a single layer of elements exists under the FRP, the element thickness may
become too coarse to model the localised shearing of the concrete layer directly under
Chapter 7 – Finite Element Modelling of Unidirectional and Bidirectional Fiber Patch
Anchors
169
the FRP plate. In order to investigate this effect, element thicknesses of 2.5mm, 5mm
and 15mm were investigated. Marginal deviations in the FE predictions from the
experimental data are observed when the finite element thickness in (y) is greater than
5mm (mesh 4). Such behaviour was expected, due to the element being unable to
capture the local stress concentrations below the FRP bond line to a sufficient level of
refinement. It is apparent that a coarser mesh size resulted in a stiffer post-peak
response and a slightly higher load prediction. The stiffer response and higher load
predictions can be explained by the model assuming a uniform crack distribution over a
relatively large finite element, while the actual cracks are concentrated in a much
smaller cracked region of the element.
(a) (b)
0
20
40
60
80
100
120
0 500 1000 1500 2000 2500
Load
(kN)
Microstrain
G3 (SG)G3 (AR)G3 (FEM) Mesh1G3 (FEM) Mesh2G3 (FEM) Mesh3G3 (FEM) Mesh4
0
20
40
60
80
100
120
0 500 1000 1500 2000 2500
Load
(kN)
Microstrain
G4 (SG)G4 (AR)G4 (FEM) Mesh1G4 (FEM) Mesh2G4 (FEM) Mesh3G4 (FEM) Mesh4
Chapter 7 – Finite Element Modelling of Unidirectional and Bidirectional Fiber Patch
Anchors
170
(c) (d)
Figure 7.13 - Comparison of load-strain curves predicted by the models with different mesh sizes
The load-slip curves are compared in Figure 7.14 at distances of 75mm and 125mm
away from the concrete free edge. Meshes 1-3 show acceptable levels of correlation.
However, the model which used courser mesh size (mesh 4) overestimated the peak
shear stress by 27% at 75mm away from the concrete free edge. The error is attributed
to the stiffer predictions obtained for this mesh size.
The 20 x 5 x 25mm element size used in mesh 3 was chosen as the basic template for
all specimens. It was found that mesh 3 produced predictions to a good level of
accuracy while enabling tasks to be easily be accomplished within the available
computational resources.
0
20
40
60
80
100
120
0 500 1000 1500 2000 2500
Load
(kN)
Microstrain
G5 (SG)G5 (AR)G3 (FEM) Mesh1G3 (FEM) Mesh2G5 (FEM) Mesh3G3 (FEM) Mesh4
0
20
40
60
80
100
120
0 500 1000 1500 2000 2500Load
(kN)
Microstrain
G6 (SG)G6 (AR)G3 (FEM) Mesh1G3 (FEM) Mesh2G6 (FEM) Mesh3G3 (FEM) Mesh4
Chapter 7 – Finite Element Modelling of Unidirectional and Bidirectional Fiber Patch
Anchors
171
(a) (b)
Figure 7.14 - Comparison of bond-slip curves predicted by the models with different mesh sizes
7.5.2 Sensitivity to fracture energy
The bond strength of FRP-to-concrete joint is largely dependent on the concrete tensile
properties. The parameters used for suitable modelling of crack formation and
propagation in concrete were concrete tensile strength and mode 1 concrete fracture
energy (GFI). The model defined concrete fracture energy as the energy needed to create
a unit area of stress free crack.
Fracture energy for concrete has been determined using a total of 7 available models
summarised in chapter 3 and the results are summarised in table 7.5. The available
models tend to exhibit a range of variability between 89.8-210.2 N/m.
0
1
2
3
4
5
6
7
0 0.1 0.2 0.3 0.4
Bond
Stress
(MPa
)
Slip (mm)
75mm (FEM) Mesh175mm (FEM) Mesh275mm (FEM) Mesh375mm (FEM) Mesh475 mm (GA)
0
1
2
3
4
5
6
0 0.1 0.2 0.3 0.4
Bond
Stress
(MPa
)
Slip (mm)
125mm (FEM) Mesh1125mm (FEM) Mesh2125mm (FEM) Mesh3125mm (FEM) Mesh4125 mm (GA)
Chapter 7 – Finite Element Modelling of Unidirectional and Bidirectional Fiber Patch
Anchors
172
Source GFI (N/m)
Bazant and Oh (1983) 89.8 CEB-FIB Model Code (1990) 124.7
van Mier (1997) 101.9 Trunk and Wittmann (1998) 210.2
Neubauer and Rostasy (1999) 140.2 Bazant and Becq-Giraudon (2002) 113.3
Dia J.G & Ueda T (2003) 191.0
Notes: Fracture energy calculations based on the following concrete properties: (ft = 4.72 MPa, f’c = 62
MPa, aggregate size = 20mm)
Table 7.5 - Concrete fracture energy variations
Parametric studies were also performed within the numerical model to investigate the
effects of 5 alternative fracture energy values (100-255 N/m) on the load vs deformation
response. As depicted in figure 7.15, the upper bound predictions tended to match
experimental data most favourably. The models which used lower facture energy
exhibited higher cracking at the FRP-to-concrete interface prior to failure. This was due
to the interfacial debond cracks being able to propagate more readily, resulting in wider
transverse cracks and a subsequent drop in stiffness and predicted failure load. The
ultimate failure load was observed to drop marginally from 93.3 kN to 90.2 kN when
the fracture energy was reduced from 255 N/m to 100 N/m (3.3% variation). However
the variation is stiffness can best be observed by examining the bond-slip predictions
along the length of the laminate presented in figure 7.16.
Chapter 7 – Finite Element Modelling of Unidirectional and Bidirectional Fiber Patch
Anchors
173
(a) (b)
(c) (d)
Figure 7.15 - Comparison of load-strain curves predicted by the models with different fracture energy
The interfacial bond stress at any given load is very much dependant on the gradient of
the force vs displacement curve. A loss of stiffness at a certain distance along the length
of the laminate would inevitably be followed by a corresponding loss of bond stress.
This is best observed in figure 7.16 where the bond stress shows very little variation
with fracture energy at a distance of 75mm from the concrete edge. However, a
0102030405060708090
100
0 500 1000 1500 2000 2500
Load
(kN)
Microstrain
G3 (SG)G3 (AR)G3 (100 N/m)G3 (125 N/m)G3 (165 N/m)G3 (210 N/m)G3 (255 N/m)
0102030405060708090
100
0 500 1000 1500 2000 2500
Load
(kN)
Microstrain
G4 (SG)G4 (AR)G4 (100 N/m)G4 (125 N/m)G4 (165 N/m)G4 (210 N/m)G4 (255 N/m)
0102030405060708090
100
0 500 1000 1500 2000 2500
Load
(kN)
Microstrain
G5 (SG)G5 (AR)G5 (100 N/m)G5 (125 N/m)G5 (165 N/m)G5 (210 N/m)G5 (255 N/m)
0102030405060708090
100
0 500 1000 1500 2000 2500
Load
(kN)
Microstrain
G6 (SG)G6 (AR)G6 (100 N/m)G6 (125 N/m)G6 (165 N/m)G6 (210 N/m)G6 (255 N/m)
Chapter 7 – Finite Element Modelling of Unidirectional and Bidirectional Fiber Patch
Anchors
174
significant reduction of bond stress is observed with reducing fracture energy when
examined at a distance of 125mm away from the concrete loaded edge. This suggests
that the increasingly rapid crack propagation, resulting from reduced fracture energy, is
more influential at a further distance away from the start of the bond line.
Figure 7.16 - Comparison of load-slip curves predicted by the models with different fracture energy
As a result of the above , a fracture energy of (GF = 210 N/m) was adopted as the final
value to be used within the concrete model based on the fracture energy model by
(Trunk and Wittmann 1998). The value was chosen due to observation that the upper
bound predictions of fracture energy resulted in a more realistic prediction of stiffness
and bond-slip behaviour.
7.5.3 Sensitivity to adhesive stiffness
In this section, the effects of varying the adhesive stiffness on the load vs strain
response are investigated. The adhesive layer is defined as the epoxy layer in between
the concrete and the FRP laminate. The stiffness of this layer (Ea) can influence the
stress transfer between the composite and concrete as well as the opening of the inclined
0
1
2
3
4
5
6
0 0.05 0.1 0.15 0.2 0.25 0.3
Bond
Stress
(MPa
)
Slip (mm)
75mm (GA)75mm (FEM) 100 N/m75mm (FEM) 125 N/m75mm (FEM) 165 N/m75mm (FEM) 210 N/m75mm (FEM) 255 N/m
0
1
2
3
4
5
6
0 0.05 0.1 0.15 0.2 0.25 0.3
Bond
Stress
(MPa
)
Slip (mm)
125mm (GA)125mm (FEM) 100 N/m125mm (FEM) 125 N/m125mm (FEM) 165 N/m125mm (FEM) 210 N/m125mm (FEM) 255 N/m
Chapter 7 – Finite Element Modelling of Unidirectional and Bidirectional Fiber Patch
Anchors
175
cracks. To study the sensitivity of the finite element model, five different values of Ea
were adopted. They are listed in Table 7.6.
Designation Adhesive elastic modulus Ea (GPa) A1 3.0 A2 5.0 A3 7.5 A4 10 A5 15
Table 7.6 - Adhesive modulus variations
The load vs slip curves for specimens with different Ea values are compared in figure
7.17. There is little difference in the behaviour of the models except for a very slight
reduction in stiffness (with reducing Ea values) in the early stages of loading up to a
load of approximately 70 kN. No significant differences in bond-slip response are
expected. The value for adhesive stiffness chosen for inclusion within the final
numerical model was 10 GPa – which was also the value quoted by the manufacturer
and verified by experimental testing.
(a) (b)
0
10
20
30
40
50
60
70
80
90
100
0 500 1000 1500 2000 2500
Load
(kN)
Microstrain
G4 (SG)G4 (AR)G4(A 3 GPa)G4(A 5 GPa)G4(A 7.5 GPa)G4(A 10 GPa)G4(A 15 GPa)
0102030405060708090
100
0 500 1000 1500 2000 2500
Load
(kN)
Microstrain
G3 (SG)G3 (AR)G3(A 3 GPa)G3(A 5 GPa)G3(A 7.5 GPa)G3(A 10 GPa)G3(A 15 GPa)
Chapter 7 – Finite Element Modelling of Unidirectional and Bidirectional Fiber Patch
Anchors
176
(c) (d)
Figure 7.17 - Comparison of load-strain curves predicted by the models with different adhesive stiffness
7.5.4 Sensitivity to concrete strength
Parametric studies based on a well calibrated numerical model can be used to expand
the existing experimental data while maintaining a minimal number of tests. Parametric
studies were conducted on anchorage type 5 to investigate the effects of concrete
strength on the maximum strain reached prior to debond. Three alternative concrete
strengths were chosen (32, 45, 62 MPa) and corresponding parameters within the
concrete and interface material models were adjusted accordingly and presented in table
7.7.
Since experimental results of varying concrete strength are not available the respective
material parameters such as the concrete young’s modulus, concrete tensile strength and
shear strengths of the interface could be calculated from the compressive strength using
the following expressions, the model for concrete shear strength (input as cohesion) was
based on (JCI 2003) :
(1) (2) (3)
0102030405060708090
100
0 500 1000 1500 2000 2500
Load
(kN)
Microstrain
G5 (SG)G5 (AR)G5(A 3 GPa)G5(A 5 GPa)G5(A 7.5 GPa)G5(A 10 GPa)G5(A 15 GPa)
0102030405060708090
100
0 500 1000 1500 2000 2500
Load
(kN)
Microstrain
G6 (SG)G6 (AR)G6(A 3 GPa)G6(A 5 GPa)G6(A 7.5 GPa)G6(A 10 GPa)G6(A 15 GPa)
Chapter 7 – Finite Element Modelling of Unidirectional and Bidirectional Fiber Patch
Anchors
177
Concrete fracture energy was adjusted accordingly for each concrete strength, while
maintaining the same model used by (Trunk and Wittmann 1998) with a reduction in
aggregate size. The load vs strain curved predicted by the models with different
concrete strengths are depicted in figure 7.18.
Parameters Case 1 Case 2 Base Model f'c (MPa) 32 45 62 Ec (MPa) 28,567 33,876 39,760 ft (MPa) 3.39 4.02 4.99 GF
I (N/m) 150 178 210 c (MPa) 4.3 5.0 5.49
,max 3632 4227 4707 P,max (kN) 161.2 187.7 209
Table 7.7 – Summary of material properties used to evaluate sensitivity to concrete strength
(a) (b)
020406080
100120140160180200220
0 1000 2000 3000 4000
Load
(kN)
Micro strain ( )
G4 (62MPa)G4 (45MPa) 0
20406080
100120140160180200220
0 10002000300040005000
Load
(kN)
Micro strain ( )
G3 (62 MPa)G3 (45 MPa)G3 (32 MPa)
Chapter 7 – Finite Element Modelling of Unidirectional and Bidirectional Fiber Patch
Anchors
178
(c) (d)
(e)
Figure 7.18 - Comparison of load-strain curves predicted by the models with different concrete strengths
An examination of figure 7.19 reveals an approximately linear relationship between the
concrete strength and the maximum laminate strain (measured at the concrete free end)
reached prior to debond. It is therefore reasonable to assume that the numerical data can
be used to extrapolate anchorage strain efficiencies for the concrete strengths within
range and that the use of higher concrete strengths will result in higher capacities.
020406080
100120140160180200220
0 1000 2000 3000 4000
Load
(kN)
Micro strain ( )
G6 (62 MPa)G6 (45 MPa)G6 (32 MPa)
020406080
100120140160180200220
0 1000 2000 3000 4000
Load
(kN)
Micro strain ( )
G5 (62 MPa)G5 (45 MPa)G5 (32 MPa)
020406080
100120140160180200220
0 1000 2000 3000
Load
(kN)
Micro strain ( )
G7 (62 MPa)
G7 (45 MPa)
G7 (32 MPa)
Chapter 7 – Finite Element Modelling of Unidirectional and Bidirectional Fiber Patch
Anchors
179
Figure 7.19 - Anchorage Type 5 parametric study – Concrete strength vs max laminate strain prior to de-bond.
7.6 Conclusion
The correct validation of FE models creates a powerful tool to investigate the results of
alternative experimental configurations, variations in materials properties. Such
variations are often costly and time consuming to investigate using experimental studies
alone. The present paper has demonstrated the effectiveness of FE models in predicting
the pre-peak and post-peak responses of FRP-to-concrete joints anchored using
unidirectional and bidirectional fibers. The numerical predictions achieved close
correlations with the experimental data. FRP fibers modelled as smeared reinforcement
within a homogenous saturant matrix was found to simulate the FRP unidirectional and
bidirectional fiber behaviours with reasonable accuracy. However, the modelling
approaches and the assumptions used to define the concrete cracking model, saturant
homogeneity and idealised bond between the adhesive and FRP can result in a higher
stiffness and have marginal influence on the results. The addition of the bidirectional
patch anchors was found to significantly enhance the strength and ductility of the FRP-
to concrete joint and was effective in providing a 93-109% increase in failure load,
which demonstrated the efficient distribution of FRP-adhesive stresses over a greater
width of concrete. Parametric studies conducted on concrete strength indicated an
approximately linear relationship between the grade of concrete and the maximum FRP
laminate strain reached prior to debond.
30
35
40
45
50
55
60
65
70
3500 4000 4500 5000Co
ncrete
streng
th(f'c)
microstrain ( )
Chapter 8 – Finite Element Investigation into the Size Effect of Bidirectional Fiber
Patch Anchors
180
8 CHAPTER 8 – FINITE ELEMENT INVESTIGATION INTO THE SIZE EFFECT OF BIDIRECTIONAL FIBER PATCH ANCHORS
8.1 Introduction
The previous chapters have shown that ±45º oriented bidirectional fabric anchorages
(herein Patch Anchors) resulted in gradual debonding of FRP laminates as a result of
FRP-adhesive stresses being distributed over a greater area of the concrete. The
anchored joints experienced increases in strength of up to 93-109 % as well as loaded
end slippage of 4 to 8 times above that of the unanchored counterparts. However, tests
conducted on these forms of anchorages have been very limited in number. Such
limitations have inspired further experimental work to investigate the parameters of
patch anchor size, spacing and fiber thickness on overall anchorage performance,
alongside further numerical simulations which will be presented in this chapter. It is
established research practice to use finite element simulations in order to reduce the
number of experimental tests, which are costly in terms of time and money. Calibrated
numerical models can be used to extend the pool of available experimental data via
parametric studies (Pham and Al-Mahaidi 2005, 2007; Al-Mahaidi and Hii 2007).
8.2 The Proposed Finite Element Model
The FE model was implemented in ATENA 3D (Cervenka 2007) and similar to the
stage 2 models, utilised axis-symmetric boundary conditions through the centre line of
the specimen to reduce model size and solution time. The components of the model
included the definition of material models for concrete, FRP laminates, FRP
bidirectional fabric, steel reinforcement and an interface bond law between the fibers
and the concrete (refer figure 8.1).
Chapter 8 – Finite Element Investigation into the Size Effect of Bidirectional Fiber
Patch Anchors
181
Figure 8.1 – Summary of finite element model components.
8.2.1 Modeling of concrete The non-linear compressive behaviour of concrete can be captured using numerical non-
linear plasticity models. Inclusion of cracking response is achieved using fracture-
plastic material models. The proposed concrete material model was based on the
smeared crack model and refined crack band theory. The compressive failure was
simulated using a biaxial stress failure criterion based on Kupfer (1969). A reduction of
compressive strength and shear stiffness after cracking was also considered by
definition of a shear factor coefficient to define the ratio between normal to shear
stiffness of cracked elements. The input parameters for the required by the concrete
material model were: Young's modulus (Es), compressive strength (f’c,), tensile strength
(fct), poisson ratio (v), mode I specific fracture energy (GFI) and shear factor coefficient.
The chosen parameters are summarised in table 8.1.
Concrete properties ModelYoung's modulus, Es (MPa) 42010 Mean compressive strength, 69.2
Characteristic tensile strength, 5.02 Poisson ratio, v 0.2
Specific fracture energy, GF 210 Shear factor coefficient 20
Table 8.1 - Concrete material model parameters used in numerical model
Full bond between
each layer of fabric
Interface Model between
concrete and 1st layer of fabric
Concret
FRP Laminate
placed in between
each layer of
fabric
2x0.86mm thick layers of
saturant with embedded
Chapter 8 – Finite Element Investigation into the Size Effect of Bidirectional Fiber
Patch Anchors
182
A fracture energy of (GF = 210 N/m) was adopted in this study was based on (Trunk
and Wittmann 1998) and used a maximum aggregate size of 20mm. In order to define
the relationship between normal and shear crack stiffness, a shear factor coefficient of
20 was specified based on experimental work by (Walraven, 1981).
8.2.2 Modeling FRP Patch Anchors
FRP bidirectional fiber sheets consisted of loosely woven fibers orientated in the ±45°
directions and embedded within a saturant matrix. The material was simplified in the FE
model by defining three dimensional brick elements of equivalent 0.86mm sheet
thickness for a single layer. This was assigned with saturant material properties which
consisted of an isotropic linear elastic material with VonMises plasticity hardening. The
fibers were defined using smeared reinforcement in perpendicular orientations (within
the saturant elements) representing the embedded fibers. The orientation of the smeared
reinforcement was defined such that it was ±45º to the direction of the laminate. A fiber
fraction of 19.7% in each direction was used to replicate the orthogonal fibers and was
obtained by dividing the total area of loose fibers by the area of saturant.
8.2.3 Modeling steel reinforcement
The steel reinforcement was assumed to be elastic, perfectly plastic and was defined
using a bi-linear stress-strain law. Steel bars were modelled individually as discrete bars
in the concrete macro elements. The assumption of perfect bond between the steel bars
and concrete elements was made relying on full strain compatibility.
8.2.4 Modeling FRP-to-Concrete Interface
The present study uses interface elements as contact between two surfaces (concrete and
FRP). A constitutive bond-slip model or shear traction-separation law was assigned to
the interface elements which were calibrated using the experimental data. This bond-slip
relationship consisted of two stages: an initially elastic stage, where the interfacial
stresses increase with the slip until the strength of the interface is reached, and a
softening stage, in which, interfacial stresses decrease with the slip eventuating in
Chapter 8 – Finite Element Investigation into the Size Effect of Bidirectional Fiber
Patch Anchors
183
debonding. Tensile and shear softening of the interface was also considered using a
multi-linear softening law calibrated with the experimental data, refer figure 8.2. The
input parameters for the interface material model are summarized in table 8.2.
The bond slip data can be computed from the axial strains of the FRP laminate
measured at discrete locations. The strain measurements obtained at gauge locations,
photogrammetry measurements and FE monitoring points can be used to obtain the
bond-slip information.
The typical interfacial bond slip curve shown in figure 8.2 was used to determine the
numerical parameters of: tangential stiffness (Ktt), cohesion (c) and friction coefficient
( ) for the shear-displacement function. The multi-linear softening component was also
derived from the experimental data. The value for cohesion input into the numerical
model (6.85 MPa) was the average of the experimental peak values obtained for all 3
control specimens. This resulted in a peak shear stress of 5.25 MPa when back
calculated from the strain values produced from the FE simulation, which also
corresponded to the average experimental shear stress value from all locations measured
across all control specimens, 5.28 MPa (refer figure 8.2 (b)).
(a) (b)
Figure 8.2 - Typical interface model behaviour in shear with cohesion softening law; (a) Numerical definition (Cervenka 2007); (b) shear-slip curve for interface derived from experimental data. Where: = interfacial shear stress, = normal stress, = friction angel, Ktt = tangential stiffness, GF
I= mode 1 fracture energy
0
1
2
3
4
5
6
7
8
9
0 0.05 0.1 0.15 0.2 0.25 0.3
Bond
Stress
(MPa
)
Slip (mm)
FEM
Exp. Fitted. Bond slip
Chapter 8 – Finite Element Investigation into the Size Effect of Bidirectional Fiber
Patch Anchors
184
Parameter Value adopted Adhesive-
to-concrete Interface
Tangential stiffness, Ktt (MN/m3) 1.5x106
Cohesion, Input into FEM model, (MPa) 6.85
Max shear stress, back calculated from FE model strains, (MPa)
5.25
Average value for max shear stress obtained from exp. data, (MPa)
5.28
Table 8.2 - Interface material model parameters used in numerical model
8.3 Results of non-linear finite element analyses
To verify that the finite element models are simulating the behaviour of the bond
specimens properly, four items are compared between the experimental and numerical
results. They are the peak load attained, crack patterns, failure mode, the slip behaviour
and the strain distribution along the length of the FRP plate.
In tables and figures which follow reference is made to V3D (Photogrammetry) and SG
(strain gauge). These refer to the two data acquisition techniques used in the
experimental program.
8.3.1 Crack patterns and failure modes
From the non-linear finite element analyses, ‘open’ crack patterns were produced as
illustrated in figure 8.3 for progressive levels of loading. In all models, the crack
patterns at or after the peak load step are shown. The width of the crack can be visually
observed by examining the thickness of the lines crossing the cracked element and the
crack direction can be determined by observing the orientation. From the plots, it is
clear that cracking initiated at the loaded face once the interfacial tensile and shear
stresses reached the strength parameters of the concrete. Cracking subsequently
progressed horizontally long the length of the laminate, which was coupled with wider
Chapter 8 – Finite Element Investigation into the Size Effect of Bidirectional Fiber
Patch Anchors
185
crack widths as the load level increased. Almost all of the concrete cracking observed
below the interface was concentrated within a 15mm depth below the FRP bond line.
(a) 50 kN (b) 77kN
(c) Failure
Figure 8.3 – Failure model of Control Specimen (FEM Model) depicting exaggerated deformations
Two modes of failure can be observed within the FEM exaggerated deformation plots
presented in figures 8.4 to 8.8 at failure. Concrete cover separation failure can be
identified by excessive cracking and deformation of the concrete elements directly
beneath the adhesive layer, whilst the bond is maintained in tack. Concrete-adhesive
interfacial debonding can be observed by a distinct separation of the adhesive or
saturant material from the concrete surface, resulting in a visible gap between the
elements. This is indicative that the strength criterion of the interface elements has been
exceeded.
Principal Tensile Stress
(MPa) Concrete cover separation failure
Chapter 8 – Finite Element Investigation into the Size Effect of Bidirectional Fiber
Patch Anchors
186
Figures 8.3 to 8.7 also depict the principle stresses recorded within a narrow band stress
range of -1.5 to 10 MPa. Although stresses within the laminate greatly exceeded this
figure the stress band was chosen to clearly capture the local stress distributions within
the concrete and patch anchor materials with progressive loading.
Based on the failure criteria described earlier, the failure modes produced by the finite
element models were similar to those observed in the experiments. Concrete cover
separation failure was the predominating mode produced by the FE analysis for the
control specimen (refer figure 8.4c) which matched closely with the experimental data.
The majority tensile stresses in the concrete (unanchored model) were concentrated
under the width of the FRP laminate and up to 50mm away from the laminate edge. The
addition of bidirectional fabric in anchorage types 1-4 clearly improved the stress
distribution by engaging a wider width of concrete. This was the main mechanism
whereby a higher anchorage strength was reached.
According to the finite element prediction for anchor type 1, concrete-adhesive
interfacial debonding dominated, which was also the predominating failure mode
observed in the experiments. The remaining anchorage types failed by either debonding
within the concrete cover zone or the adhesive-to-concrete interface.
It should be noted that laminate slippage was not an observed failure mode in the finite
element simulations. Since laminate slippage and patch anchor debond were found to
occur at similar load levels within the experiments, the lack of replicating this type of
failure did not interfere with the overall peak strength values of the joints.
Chapter 8 – Finite Element Investigation into the Size Effect of Bidirectional Fiber
Patch Anchors
187
(b) 50 kN (b) 100 kN
(c)150 kN
(e) Failure
Figure 8.4 – Failure model of Anchor Type 1 (FEM Model) depicting exaggerated deformations
Principal Tensile Stress
(MPa)
Concrete-Adhesive
Interfacial Debonding
Chapter 8 – Finite Element Investigation into the Size Effect of Bidirectional Fiber
Patch Anchors
188
(a) 50 kN (b) 100 kN
(c)138 kN
(e) Failure
Figure 8.5 – Failure model of Anchor Type 2 (FEM Model) depicting exaggerated deformations
Concrete cover separation failure
Principal Tensile Stress
(MPa)
Chapter 8 – Finite Element Investigation into the Size Effect of Bidirectional Fiber
Patch Anchors
189
(a) 50 kN (b) 100 kN
(c)150 kN
(e) Failure
Figure 8.6 – Failure model of Anchor Type 3 (FEM Model) depicting exaggerated deformations
Concrete-Adhesive Interfacial
Concrete Cover
separation failure
Principal Tensile Stress
(MPa)
Chapter 8 – Finite Element Investigation into the Size Effect of Bidirectional Fiber
Patch Anchors
190
(a) 50 kN (b) 100 kN
(c)117 kN
(e) Failure
Figure 8.7 – Failure model of Anchor Type 4 (FEM Model) depicting exaggerated deformations
Concrete-Adhesive
Interfacial Debonding
Concrete Cover
separation failure
Concrete Cover
separation failure
Principal Tensile Stress
(MPa)
Chapter 8 – Finite Element Investigation into the Size Effect of Bidirectional Fiber
Patch Anchors
191
8.3.2 Peak loads
The peak loads obtained from the experiment and numerical models are summarised in
Table 8.3. For the experiment, the peak loads shown in the table are the total loads
measured. For the numerical model, the peak loads are multiplied by a factor of two
since only half of the specimen was modelled.
Specimen
Width of
Patch
Anchor
(mm)
Length
of
Patch
Anchor
Exp
Failure
Load
(kN)
Exp Max
Laminate
strain
( )
FEM
Failure
Load
(kN)
FEM
Accuracy
(±%)
0.1 NA NA 83.8 2875 78.1 -6.9 0.2 NA NA 81.0 3062 78.1 -3.7 0.3 NA NA 83.0 3100 78.1 -6.1 1.1 400 300 131.0 4406 151.8 +15.9 1.2 400 300 140.2 4922 151.8 +8.3 2.1 400 250 111.0 3819 138.9 +25.1 2.2 400 250 128.1 4328 138.9 +8.4 3.1 300 300 151.6 5378 150.1 -1.1 3.2 300 300 138.5 4801 150.1 +8.3 3.3 300 300 158.8 5600 150.1 -5.5 3.4 300 300 139.1 5091 150.1 +7.8 4.1 200 300 140.6 4950 117.1 -16.8 4.2 200 300 119.9 4504 117.1 -2.4 4.3 200 300 112.5 4124 117.1 +4.0 4.4 200 300 123.6 4514 117.1 -5.3
Table 8.3 – Results summary
It is apparent that control specimens 0.1-0.3 exhibited consistent failure loads and
elongations prior to debond with only 3.4% variation. All anchorage specimens 1-4,
exhibited improvements in strength and deformation compared with their unanchored
counterparts. Since specimens 2.1 and 2.2 were designed with a lower patch anchorage
length (250mm), they tended to exhibit laminate slippage at a lower load rather than
patch anchor debond. Specimens with anchorage lengths of 300mm failed by either
laminate slippage at a higher load or patch anchor debond. Due to the tendency for
premature slippage to occur when using lower anchorage lengths it was decided that a
Chapter 8 – Finite Element Investigation into the Size Effect of Bidirectional Fiber
Patch Anchors
192
minimum length of 300mm be used for future study. Specimens which used 300mm
anchor lengths and failed by laminate slippage at load levels prior to failure between
138.5-151.6 kN (9.5% variation). Patch anchor debond was observed for the same
anchorage length at similarly high levels, 131.0-140.2 kN for anchor widths of 320mm
and greater and at a lower load of 119.9 kN for 220mm wide anchors.
8.3.3 FRP strain distributions along length of laminate
The experimental and numerical strain distributions along the bonded joint at selected
load levels are compared in figures 8.8 to 8.12. At each load level, the experimental and
numerical load levels were selected to be as close as possible to each other. For the
numerical results, each discrete data point corresponds to the average strain calculated
at the location of the nodes at the top of the FRP plates.
The load vs strain curves are depicted at gauge location G2 to G7 corresponding to a
distance of 280mm along the strain vs distance curve. The edge of the concrete block
corresponds to a distance along the laminate of 150mm.
The FE model tended to predict a slightly lower failure load for the control specimens
within 3.7-6.9% accuracy. Good correlations in stiffness between the FE and
experimental results can be observed for strain gauges G2, G5-G7. However stiffer
predictions are observed for gauges G3 and G4 up to a load level of 65 kN. The FE
model also predicts a more gradual debonding failure which can be observed by the
gradual reduction in gradient of the load vs strain curves as opposed to the experimental
results where the energy release as a result of debonding was more sudden. This was not
an observation specific to the control specimen, but a general trend observed across all
models.
Anchorage specimens 1-3 showed similar predictions in stiffness between the FE and
experimental values and the shapes and distributions of the curves are very similar. The
FE results tended to predict higher failure loads for anchorage types 1-3, which are seen
in table 8.3. The results indicate that the numerical models overestimated the failure
load of specimen type 1 by an average value of +11.9%, type 2 by +16.2%, type 3 by
+2.0 % and underestimated type 4 by -5.76%. The deviations from the actual behaviour
Chapter 8 – Finite Element Investigation into the Size Effect of Bidirectional Fiber
Patch Anchors
193
can be explained by examining the assumptions inherent within the numerical model.
For instance, the assumptions inherent within the concrete model (as described by the
smeared crack band approach) fail to consider discrete cracking, which can result in
local stress concentrations within the bond line. In practice, discrete cracking may
initiate debonding at a lower load due to local stress concentrations inducing peeling
stresses at the crack location. In addition, modelling the bidirectional fabric as smeared
reinforcement within a uniform saturant matrix resulted in a slightly stiffer material and
a more efficient distribution of stress through the patch anchor to a wider area of
concrete. It is not possible to model the non-uniformities that the fibers create within the
saturant matrix and imperfections in saturant homogeneity. Lastly, the assumption of
idealised bond between adhesive and concrete layers – adhering to a pre-defined
interface bond law, ignores local imperfections in the bond line which can result from
application. All of the above may contribute to the higher failure load predictions.
However, despite the various assumptions required in the FE model, it is evident that
the predictions are reasonably close to the experimental values.
Chapter 8 – Finite Element Investigation into the Size Effect of Bidirectional Fiber
Patch Anchors
194
(a)
(b) (c) (d)
(e) (f) (g) Figure 8.8 – Type 0.1 Strain distribution Summary: (a) Strain vs Distance along FRP laminate; (b) Gauge G2; (c) Gauge G3; (d) Gauge G4; (e) Gauge G5; (f) Gauge G6 ; (g) Gauge G7
0
500
1000
1500
2000
2500
3000
0 100 200 300 400 500
Micro
strain
()
Distance along Laminate from gauge G1 (mm)
50kN (FEM)
30kN (FEM)
40kN (FEM)
60kN (FEM)
70kN (FEM)
77.38kN (FEM)
30kN (SG)
40kN (SG)
50kN (SG)
60kN (SG)
70kN (SG)
80kN (SG)
84kN (SG)
0
10
20
30
40
50
60
70
80
90
0 1000 2000 3000
Load
(kN)
Micro strain ( )
G2 (FEM)
G2 (SG)
G2 (V3D)
0
10
20
30
40
50
60
70
80
90
0 1000 2000 3000
Load
(kN)
Micro strain ( )
G3 (FEM)
G3 (SG)
G3 (V3D)
0
10
20
30
40
50
60
70
80
90
0 1000 2000 3000
Load
(kN)
Micro strain ( )
G4 (FEM)
G4 (SG)
G4 (V3D)
0
10
20
30
40
50
60
70
80
90
0 1000 2000 3000
Load
(kN)
Micro strain ( )
G5 (FEM)
G5 (SG)
G5 (V3D)
0
10
20
30
40
50
60
70
80
90
0 500 1000 1500 2000
Load
(kN)
Micro strain ( )
G6 (FEM)
G6 (SG)
G6 (V3D)
0
10
20
30
40
50
60
70
80
90
0 500 1000 1500
Load
(kN)
Micro strain ( )
G7 (FEM)
G7 (SG)
G7 (V3D)
Chapter 8 – Finite Element Investigation into the Size Effect of Bidirectional Fiber
Patch Anchors
195
(a)
(b) (c) (d)
(e) (f) (g) Figure 8.9 – Type 1.2 Strain distribution Summary: (a) Strain vs Distance along FRP laminate; (b) Gauge G2; (c) Gauge G3; (d) Gauge G4; (e) Gauge G5; (f) Gauge G6 ; (g) Gauge G7
0
500
1000
1500
2000
2500
3000
3500
4000
4500
5000
5500
0 100 200 300 400
Micro
strain
()
Distance along Laminate from gauge G1 (mm)
20kN (SG)
40kN (SG)
60kN (SG)
80kN (SG)
100kN (SG)
140kN (SG)
20kN (FEM)
40kN (FEM)
60kN (FEM)
80kN (FEM)
100kN (FEM)
151.9kN (FEM)
0
20
40
60
80
100
120
140
160
0 1000 2000 3000 4000 5000
Load
(kN)
Micro strain ( )
G2 (FEM)
G2 (SG)
G2 (V3D)
0
20
40
60
80
100
120
140
160
0 1000 2000 3000 4000 5000
Load
(kN)
Micro strain ( )
G3 (FEM)G3 (SG)G3 (V3D)
0
20
40
60
80
100
120
140
160
0 1000 2000 3000 4000
Load
(kN)
Micro strain ( )
G4 (FEM)
G4 (SG)
G4 (V3D)
0
20
40
60
80
100
120
140
160
0 1000 2000 3000 4000
Load
(kN)
Micro strain ( )
G5 (FEM)
G5 (SG)
G5 (V3D)
0
20
40
60
80
100
120
140
160
0 1000 2000 3000
Load
(kN)
Micro strain ( )
G6 (FEM)G6 (SG)G6 (V3D)
0
20
40
60
80
100
120
140
160
0 500 1000 1500 2000
Load
(kN)
Micro strain ( )
G7 (FEM)G7 (SG)G7 (V3D)
Chapter 8 – Finite Element Investigation into the Size Effect of Bidirectional Fiber
Patch Anchors
196
(a)
(b) (c) (d)
(e) (f) (g) Figure 8.10 – Type 2.2 Strain distribution Summary: (a) Strain vs Distance along FRP laminate; (b) Gauge G2; (c) Gauge G3; (d) Gauge G4; (e) Gauge G5; (f) Gauge G6 ; (g) Gauge G7.
0
500
1000
1500
2000
2500
3000
3500
4000
4500
5000
5500
0 100 200 300 400
Micro
strain
()
Distance along Laminate from gauge G1 (mm)
10kN (FEM)
20kN (FEM)
40kN (FEM)
60kN (FEM)
100kN (FEM)
120kN (FEM)
144kN (FEM)
20kN (SG)
40kN (SG)
60kN (SG)
80kN (SG)
100kN (SG)
120kN (SG)
127.9kN (SG)
80kN (FEM)
0
20
40
60
80
100
120
140
160
0 1000 2000 3000 4000 5000
Load
(kN)
Micro strain ( )
G2 (FEM)G2 (SG)G2 (V3D)
0
20
40
60
80
100
120
140
160
0 1000 2000 3000 4000
Load
(kN)
Micro strain ( )
G3 (FEM)G3 (SG)G3 (V3D)
0
20
40
60
80
100
120
140
160
0 1000 2000 3000 4000
Load
(kN)
Micro strain ( )
G4 (FEM)G4 (SG)G4 (V3D)
0
20
40
60
80
100
120
140
160
0 1000 2000 3000
Load
(kN)
Micro strain ( )
G5 (FEM)G5 (SG)G5 (V3D)
0
20
40
60
80
100
120
140
160
0 500 1000 1500 2000 2500
Load
(kN)
Micro strain ( )
G6 (FEM)G6 (SG)G6 (V3D)
0
20
40
60
80
100
120
140
160
0 500 1000 1500
Load
(kN)
Micro strain ( )
G7 (FEM)G7 (SG)G7 (V3D)
Chapter 8 – Finite Element Investigation into the Size Effect of Bidirectional Fiber
Patch Anchors
197
(a)
(b) (c) (d)
(e) (f) (g) Figure 8.11 – Type 3.4 Strain distribution Summary: (a) Strain vs Distance along FRP laminate; (b) Gauge G2; (c) Gauge G3; (d) Gauge G4; (e) Gauge G5; (f) Gauge G6 ; (g) Gauge G7.
0
500
1000
1500
2000
2500
3000
3500
4000
4500
5000
5500
0 100 200 300 400
Micro
strain
()
Distance along Laminate from gauge G1 (mm)
20kN (SG)
40kN (SG)
60kN (SG)
80kN (SG)
100kN (SG)
130kN (SG)
138.8kN (SG)
20kN (FEM)
40kN (FEM)
60kN (FEM)
80kN (FEM)
100kN (FEM)
130kN (FEM)
150kN (FEM)
0
20
40
60
80
100
120
140
160
0 1000 2000 3000 4000 5000
Load
(kN)
Micro strain ( )
G2 (FEM)G2 (SG)G2 (V3D)
0
20
40
60
80
100
120
140
160
0 1000 2000 3000 4000 5000
Load
(kN)
Micro strain ( )
G3 (FEM)G3 (SG)G3 (V3D)
0
20
40
60
80
100
120
140
160
0 1000 2000 3000 4000
Load
(kN)
Micro strain ( )
G4 (FEM)
G4 (SG)
G4 (V3D)
0
20
40
60
80
100
120
140
160
0 1000 2000 3000 4000 5000
Load
(kN)
Micro strain ( )
G5 (FEM)G5 (SG)G5 (V3D)
0
20
40
60
80
100
120
140
160
0 500 10001500200025003000
Load
(kN)
Micro strain ( )
G6 (FEM)G6 (SG)G6 (V3D)
0
20
40
60
80
100
120
140
160
0 500 1000 1500 2000
Load
(kN)
Micro strain ( )
G7 (FEM)G7 (SG)G7 (V3D)
Chapter 8 – Finite Element Investigation into the Size Effect of Bidirectional Fiber
Patch Anchors
198
(a)
(b) (c) (d)
(e) (f) (g) Figure 8.12 – Type 4.4 Strain distribution Summary: (a) Strain vs Distance along FRP laminate; (b) Gauge G2; (c) Gauge G3; (d) Gauge G4; (e) Gauge G5; (f) Gauge G6 ; (g) Gauge G7.
0
500
1000
1500
2000
2500
3000
3500
4000
4500
0 100 200 300 400
Micro
strain
()
Distance along Laminate from gauge G1 (mm)
20kN (SG)
40kN (SG)
60kN (SG)
80kN (SG)
100kN (SG)
110kN (SG)
123kN (SG)
20kN (FEM)
40kN (FEM)
60kN (FEM)
80kN (FEM)
100kN (FEM)
110kN (FEM)
117kN (FEM)
0
20
40
60
80
100
120
140
0 1000 2000 3000 4000 5000
Load
(kN)
Micro strain ( )
G2 (FEM)
G2 (SG)
G2 (V3D)
0
20
40
60
80
100
120
140
0 1000 2000 3000 4000
Load
(kN)
Micro strain ( )
G3 (FEM)G3 (SG)G3 (V3D)
0
20
40
60
80
100
120
140
0 1000 2000 3000
Load
(kN)
Micro strain ( )
G4 (FEM)G4 (SG)G4 (V3D)
0
20
40
60
80
100
120
140
0 500 1000 1500 2000
Load
(kN)
Micro strain ( )
G5 (FEM)
G5 (SG)
G5 (V3D)
0
20
40
60
80
100
120
140
0 500 1000 1500 2000
Load
(kN)
Micro strain ( )
G6 (FEM)
G6 (SG)
G6 (V3D)
0
20
40
60
80
100
120
140
0 100 200 300 400 500
Load
(kN)
Micro strain ( )
G7 (FEM)
G7 (SG)
G7 (V3D)
Chapter 8 – Finite Element Investigation into the Size Effect of Bidirectional Fiber
Patch Anchors
199
8.3.4 Strain in Bidirectional fibers Strain measurements of the bi-directions fibers at various stages of loading were
obtained from both surface mounted strain gauges and image correlation
photogrammetry. The data acquired from the experiments was then compared with the
predicted values obtained from the numerical model and the results are summarized in
figure 8.13. The strains depicted are orientated at ±45 degree angles from the
longitudinal axis and correspond to the principal direction of the bidirectional fibers
either side of the laminate. The numerical predictions generally show good correlation
with the measured values. However the FE results overestimated the strains in the
bidirectional fibers at 100mm either side of the laminate centerline for specimen 4.4. In
all cases the strains in the bidirectional fibers are concentrated within the initial 50mm
away from the laminate edge, which provides insight into the minimum spacing that
patch anchored FRP laminates may be placed beside one another without exceeding the
strength of the bond-line due to superposition of stresses. As a result, it is recommended
that where laminates are to be placed side by side along with continuous anchorage –
laminates should not be spaced closer than 250mm center-to-center, without incurring
losses in anchorage strength.
(a) (b)
0500
10001500200025003000350040004500
200 150 100 50 0 50 100 150 200
Microstrain
()
Distance from across concrete block from centreof laminate (mm)
120 kN (FEM) 120kN (SG)120kN (SG) 120kN (V3D)120kN (V3D) 120 kN (FEM)
0500
100015002000250030003500400045005000
200 150 100 50 0 50 100 150 200
Microstrain
()
Distance from across concrete block fromcentre of laminate (mm)
128kN (SG) 128kN (SG)128kN (V3D) 128kN (V3D)128 kN (FEM) 128 kN (FEM)
Chapter 8 – Finite Element Investigation into the Size Effect of Bidirectional Fiber
Patch Anchors
200
(c) (d) Figure 8.13 – Strain of 45º Bidirectional FRP either side of laminate; (a) Spec 1.2, (b) Spec 2.2, (c) Spec 3.4, (d) Spec 4.4
8.4 Parametric studies
8.4.1 Sensitivity to concrete strength
In this section, the effects of varying the concrete strength on the peak load and load-
strain distribution were investigated. Numerical models with concrete strengths of 32
and 45 MPa were analysed and compared against the base model of 69.2 MPa. The
material properties used in each respective model are summarised in table 8.4.
The shear strength of the interface was adopted from the model proposed by (JCI 2003)
and the fracture energy of the concrete was determined using the previously adopted
model by (Trunk and Wittmann 1998). However, the aggregate size was adjusted in
accordance with the concrete strength, resulting in 7mm diameter aggregates used for
32 MPa and 12mm for 45 MPa. The respective aggregate sizes were used to calculate
fracture energies values of 150 N/m and 178 N/m for 32 MPa and 45 MPa concrete
strengths. The determination of ‘failure’ for the models was based on the criterion
described earlier. In the base model for anchor type 3, ‘failure’ was found to occur by
separation of the interface elements between the adhesive and concrete. The same
failure mode was observed for both specimens of lower concrete strength.
0
1000
2000
3000
4000
5000
200 150 100 50 0 50 100 150 200
Microstrain
()
Distance from across concrete block from centreof laminate (mm)
138kN (SG) 138kN (SG)138kN (V3D) 138kN (V3D)138 KN (FEM) 138 KN (FEM)
0
1000
2000
3000
4000
200 150 100 50 0 50 100 150 200
Microstrain
()
Distance from across concrete block fromcentre of laminate (mm)
123kN (V3D) 123kN (V3D)123kN (SG) 123kN (SG)128 kN (FEM) 128 kN (FEM)
Chapter 8 – Finite Element Investigation into the Size Effect of Bidirectional Fiber
Patch Anchors
201
Parameters Case 1 Case 2 Base Model f'c (MPa) 32 45 69.2 Ec (MPa) 28,567 33,876 42009 ft (MPa) 3.39 4.02 5.02 GF
I (N/m) 150 178 210 c, (MPa) 4.3 5.0 6.85
,max 3632 4227 5106 P,max (kN) 161.2 187.7 150.1
Table 8.4 – Summary of material properties used to evaluate sensitivity to concrete strength
The load strain curves for strain gauges G3 to G7 are depicted in figure 8.14. The main
observation is that the higher concrete strength increases both the strength and stiffness
of the anchor. Further conclusions can be drawn by examining figure 8.15 which plots
the maximum strains reached against the three respective concrete strengths. The figure
reinforces the notion of a linear relationship existing between the strength of the anchor
and the compressive strength of the concrete.
(a) (b)
0
20
40
60
80
100
120
140
160
0 1000 2000 3000 4000 5000
Load
(kN)
Micro strain ( )
G3 (32 MPa)
G3 (45 MPa)
G3 (69 MPa)
0
20
40
60
80
100
120
140
160
0 1000 2000 3000 4000 5000
Load
(kN)
Micro strain ( )
G4 (32 MPa)
G4 (45 MPa)
G4 (69 MPa)
Chapter 8 – Finite Element Investigation into the Size Effect of Bidirectional Fiber
Patch Anchors
202
(c) (d)
(e)
Figure 8.14 - Comparison of load-strain curves predicted by the models with different concrete strengths
Figure 8.15 - Anchorage Type 5 parametric study – Concrete strength vs max laminate strain prior to de-bond.
0
20
40
60
80
100
120
140
160
0 1000 2000 3000 4000 5000
Load
(kN)
Micro strain ( )
G5 (32 MPa)
G5 (45 MPa)
G5 (69 MPa)
0
20
40
60
80
100
120
140
160
0 500 1000 1500 2000 2500 3000
Load
(kN)
Micro strain ( )
G6 (32 MPa)
G6 (45 MPa)
G6 (69 MPa)
0
20
40
60
80
100
120
140
160
0 500 1000 1500 2000
Load
(kN)
Micro strain ( )
G7 (32 MPa)
G7 (45 MPa)
G7 (69 MPa)
202530354045505560657075
3500 4000 4500 5000 5500
Concrete
Streng
th(M
Pa))
Micro strain ( )
Chapter 8 – Finite Element Investigation into the Size Effect of Bidirectional Fiber
Patch Anchors
203
8.5 Summary
The present study focused on the development of a numerical FE model to simulate the
behaviour of FRP patch anchors, which were capable of increasing anchorage strength
by 53-81%. The calibration of the constituent material models is described in some
detail and the parameters defining the adhesive-to-concrete interface law are presented.
The overall numerical predictions were achieved in close correlation with the
experimental data. The key findings allow the following conclusions to be made:
Non-liner Finite Element Numerical models (FEM) when correctly validated by
experimental data is capable of predicting interfacial bond properties between
concrete and FRP.
Experimental bond-slip relations can be used to inform the definition of
numerical bond-interface laws between the adhesive and FRP materials.
FRP fibers modelled as smeared reinforcement within a homogenous,
orthotropic, linear-elastic base material, designated with VonMises plasticity
hardening can model FRP fabric material behaviour to a reasonable level of
accuracy.
Patch anchorage lengths of 250mm exhibited slippage at a lower load. As a
result, it was recommended that a minimum anchorage length of 300mm be used
for future studies.
By examining the strain distributions within the bidirectional fibers it is
expected that laminates can be spaced as closely a 250mm center-to center
without any reductions in anchorage strength.
Chapter 9 – Development of Patch Anchor Prediction Model
204
CHAPTER 9 – DEVELOPMENT OF PATCH ANCHOR 9
PREDICTION MODEL 9
9.1 Introduction
The main obstacle presently preventing the widespread use of FRP anchorage measures
is that no rational and reliable design rules currently exist. As a result, FRP design
guidelines stipulate that the practical implementation of anchorage devices should be
substantiated by representative experimental testing (ACI 440.2R-08 2008). However, it
does not specify types of testing procedures that are considered adequate (Grelle and
Sneed 2011). The repercussions of time and budget constraints on small and large scale
industrial projects means that such testing is rarely carried out in practice. As a result,
the potential benefits of FRP anchorages have typically been superseded by more
conservative strengthening approaches such as section enlargement or column insertion.
Based on this study, the use of bidirectional fiber patch anchorages has demonstrated to
be a highly effective form of anchorage with the potential for applications to a wide
variety of strengthening projects. The tests conducted to date have provided promising
results, and could be utilised directly provided that the materials matched those used in
the experimental study. However enough data has been collected to attempt the
development of a prediction model which could empirically relate parameters such as
concrete strength, laminate thickness, width and spacing and patch anchor size. Such a
model would be immensely useful to both, researchers and designers.
9.2 Assessment of prediction models
In Chapter 3, several anchorage strength models available in literature were
summarised. It was found that there are a number of models derived from experimental
data, fracture mechanics and design proposals to predict the load capacity of a joint
between FRP and concrete.
These models are based on different experimental data and/or different theoretical
assumptions. Therefore, there is a need to investigate the accuracy of the models. In this
Chapter 9 – Development of Patch Anchor Prediction Model
205
section, existing prediction models are assessed and compared with the results obtained
from the experimental data. Table 9.1 summarises the strength predictions obtained for
the control specimens in stages 1 and 2 using several models proposed by researchers.
The models are observed to exhibit a wide range of variability in the predicted failure
load with values ranging from 77.1 kN to 147.7 kN for stage 1. While Chaallal 1998;
Chen and Teng 2001; and Khalifa et al. 1998 provided the closest strength predictions
to the experimental values for stages 1 and 2 of the experiments, the majority of the
proposed models overestimated the failure load. Only two models provided strength
predictions lower than the actual value for stage 2 (Maeda 1997; Yang 2001).
Model Predicted Failure
Load kN (Stage 1)
COV %
Predicted Failure
Load kN (Stage 2)
COV %
(Van Gemert 1980). 141.6 42.2 125 51.3(Neubauer 1997) 127.2 27.7 97.4 17.9 (Taljsten 1994) 134.3 34.8 104 25.9 (Yuan and Wu 1999) 136.3 36.8 105.1 27.2 (Maeda et al. 1997) 77.1 -22.6 55.3 -33.1 (Khalifa et al. 1998) 115.7 16.2 91.1 10.3 (Chaallal 1998) 102.9 3.3 64 -22.5 (Chen and Teng 2001) 100 0.4 79 -4.4 (JCI 2003) 147.7 48.3 105.4 27.6 (Yang 2001) 81.7 -18.0 61 -26.2 Actual Failure load 99.6 82.1
Table 9.1 - Summary of strength prediction models compared with FRP-to-Concrete joints
9.3 Parameters influencing an anchorage prediction model
9.3.1 Concrete Strength
The concrete strength is the primary characteristic which governs the strength of the
concrete substrate to which the FRP material is bonded. As a result, the likeliness of
patch anchor debond is largely dependant of the tensile and shear strength properties of
the concrete substrate to which the FRP is bonded which, in turn, can be correlated with
Chapter 9 – Development of Patch Anchor Prediction Model
206
the concrete compressive strength. The shear strength of the FRP-to-concrete interface
could be calculated from the experimental data by monitoring the force difference
between 2 strain gauges (along the bond line) divided by the distance between the
gauges. While this approach is suitable when experimental data is readily available – a
different approach is needed when developing a numerical model expected to provide
strength predictions where no experimental data exists.
In general, researchers have found that the maximum shear strength of the concrete
substrate can be correlated to the concrete compressive strength and the FRP effective
bond length – which is dependant on the FRP modulus and thickness (Hiroyuki 1997).
Parametric studies into alternative concrete strengths (32, 45 62 and 69.2 MPa) were
performed in FE simulations which demonstrated an approximately linear relationship
between the concrete compressive strength and the maximum FRP strain reached prior
to debond. The results were obtained by varying concrete properties alone (f'c, Ec, ft,
GFI, while keeping all other parameters constant (Kalfat R and Al-Mahaidi R 2013).
The maximum shear strength of the interface used in the FE simulations for varying
concrete strengths was determined using the model proposed by (JCI 2003) – which
considers the effect of concrete compressive strength, however ignores the influence of
effective anchorage length, which has been proven to affect the peak interfacial shear
strength and cohesion reached prior to debonding.
Figure 9.1 – Summary of parametric study results conducted on concrete strength and the maximum FRP strain reached prior to debond.
f = 32.423(f'c) + 2727.1
3500
3700
3900
4100
4300
4500
4700
4900
5100
5300
30 35 40 45 50 55 60 65 70 75
microstrain(
)
Concrete strength (f'c)
Chapter 9 – Development of Patch Anchor Prediction Model
207
From figure 9.1, it is apparent that a linear relationship exists between the concrete
strength and the maximum strain reached prior to debond which can be approximated
using equation 9.1.
if 32 < < 69.2MPa (9.1)
The data used to produce figure 9.1 was based on stage 1 of the experimental program
which was based on concrete parameters shown in table 7.7, where 62 MPa was the
concrete strength used in the experimental program.
A coefficient (r1) can be applied to the maximum FRP strain reached for 62 MPa
concrete and used as a benchmark to derive the FRP strains for other concrete strength
values:
if 32 < < 69.2 MPa (9.2)
In order to account for the influence of parameters such as: effective bond length, FRP
modulus and thickness on the maximum shear strength of the concrete interface, the
model proposed by (Tanaka 1996) was modified to account for variability in concrete
strength:
(9.3)
Where the effective bond length was based on the model proposed by (JCI 2003):
: ; And (9.4)
A limitation of 2mm has been placed on the maximum laminate thickness
recommended for use with the proposed bidirectional fiber patch anchor due to an
absence of experimental data and to avoid potential laminate slippage failure.
9.3.2 FRP width
The experimental results obtained in this study indicated two possible failure modes
whereby the FRP laminate may separate from the concrete: (1) patch anchor debond and
Chapter 9 – Development of Patch Anchor Prediction Model
208
(2) laminate slippage. Of the two failure modes observed, patch anchor debond is
preferred and was found to occur at a higher load. The load level governing laminate
slippage was largely a function of the contact area between the laminate and the
bidirectional fabric which was dependant on the laminate width and the effective
anchorage length. Since laminate slippage was not an observed failure mode in stage 1
of the experiments, which used a laminate width of 120mm, a reduction coefficient (r2)
can be applied for laminate widths less than 120mm:
(9.5)
9.3.3 FRP spacing
The strains distribution in the bidirectional fibers can provide insight into the potential
stress-strain interactions and reductions in strength, due to overlapping of strain profiles
where laminates are placed in close proximity of each other under sustained load. In
general, the strain distributions within the bidirectional fibers were localised within the
initial 100mm from the laminate edge. However, where patch anchor debond was the
predominating failure mode, strains were observed to be distributed as far as 150mm
away from the laminate edge. Specimen 1.1 was used as a benchmark to provide a worst
case scenario. The specimen failed by patch anchor debonding, thereby causing the
bidirectional fiber strains to reach approximately 3500 , over a diastance of 150mm
away from the laminate edge. The resulting strain distribution is expected to cause the
greatest potential for strain interaction resulting from superposition of principal stains in
the bidirectional fiber sheet between two adjacent laminates and is depicted in figure
9.2.
Chapter 9 – Development of Patch Anchor Prediction Model
209
Figure 9.2 – Typical strain overlay in bidirectional fibers resulting from superposition of strain between two adjacent laminates.
Based on the strain distributions shown in figure 9.2, it is apparent that a 250mm
laminate spacing would not result in a sufficient stress-strain interaction to shift the
superimposed strain distribution above the peak values. However, a laminate spacing
less than 250mm would immediately result in a reduction in strength. Examining the
experimental results for the specimens which used a laminate spacing less than 250mm
confirmed the reduction in strength, which was also confirmed in the FE simulations.
To account for the strength reduction incurred where the distance between laminates is
closer than 250mm, a strength reduction coefficient (r3) is introduced based on the
reductions in strength observed in stage 2 of the experimental program, between
specimen types 3 and 4. However, in the absence of further experimental data, FRP
laminates should be spaced no closer than 200 mm centre to centre.
(9.6)
9.3.4 FRP thickness
The FRP thickness and modulus directly govern the bond stresses generated within the
FRP bond line at any given level of fiber strain. As a result, increasing the fiber
thickness or modulus will generally reduce the FRP strain required to achieve the peak
0500
100015002000250030003500400045005000
50 0 50 100 150 200 250 300
microstrain
(
distance (mm)
Principal (±45º) fiber Stain(FRP Laminate No. 1)
Principal (±45º) fiber Strain(FRP Laminate No. 2)
Superimposed Distributionof Principal Fiber Strains
Chapter 9 – Development of Patch Anchor Prediction Model
210
bond strength of the interface. The relationship is best depicted in equation 5 where it is
shown that these properties are inversely proportional to the FRP strain required to
cause debonding.
Based on a number of experimental studies, researchers have discovered that a non-
linear relationship exists between the FRP thickness, modulus and the FRP effective
bond length – such that increasing the FRP thickness or modulus was found to increase
the effective bond length (Sato et al. 1997; Chen and Teng 2001; JCI 2003). This
phenomenon was taken into account in equations 9.8 and 9.9.
9.3.5 Anchorage length
Patch anchor lengths ranging from 250 to 300mm were investigated in stage 2 of the
experimental program presented in chapter 6. Of the two anchorage lengths
investigated, the use of 250mm long patch anchors was found to result in laminate
slippage at a lower load, which was caused by a reduction in available laminate to fabric
bond area. The overall reduction in anchorage strength, resulting from a lower patch
anchor length (250mm) was found to be approximately proportional to the ratio
between the reduced anchorage length (250mm) and the effective patch anchor length,
nominated as (300mm). In order to account for the reduction in strength ensuring from
the use of patch anchor lengths less than 300mm a further reduction factor (r4) is
introduced:
9.4 Proposed anchorage strength model
The majority of FRP bond strength models proposed by researchers calculate the pull-
off strength by multiplying the bond strength of the interface ( u) by the fiber width (bf)
and the effective bond length (Le) (Tanaka 1996; Hiroyuki 1997; Maeda et al. 1997;
Sato et al. 1997; Khalifa et al. 1998). Other researchers have proposed models which are
variations of this basic theme (Van Gemert 1980; JCI 2003).
Chapter 9 – Development of Patch Anchor Prediction Model
211
The proposed anchorage strength model uses the same basic constituent relationship
between the various influencing parameters with modification factors (r1 to r4) to
account for effects of varying concrete strength, FRP width, FRP spacing, FRP
thickness and modulus and patch anchor length. An additional factor of 1.25 is applied
to the bond strength formulations to account for the additional bond area provided by
the patch anchors. The factor was determined based on model calibrations with stage 1
of the experimental data. The model is therefore empirically derived. As a result, the
model was calibrated with the properties derived from the stage 1 experiments and
appropriate adjustment factors were applied to account for varying material properties
and anchorage sizes. The resulting expressions are summarised in equations 9.8 and 9.9.
(9.8) (9.9)
9.5 Verification of the proposed model
To verify that the proposed theoretical model is simulating the bond behaviour of the
various patch anchor configurations correctly, load and strain predictions for all
specimens tested in the experimental programs were calculated, tabulated and compared
with the actual values. For further verification, the theoretical model was also used to
provide predictions for the three alternative concrete strengths investigated in the
parametric studies conducted in the finite element simulations. The results are depicted
in table 9.2 which compares the experimental and finite element results with the
proposed model predictions.
The model was found to reasonably predict the general maximum anchorage strengths
and strains achieved prior to debond within an average accuracy of -7.8% and -5.2%
across all specimens.
Since stage 1 specimens used a patch anchor length of 270mm, the factor k4 was
reduced to 0.9 to account for the potential for laminate slippage. As a result, the model
predictions tended to be lower -10 to -20% lower than the experimental values.
However, no laminate slippage was observed in the experiments which highlights the
fact that the increased laminate width used in stage 1 (120mm) can offset the likelihood
Chapter 9 – Development of Patch Anchor Prediction Model
212
of laminate slippage when an anchorage length less than 300mm is adopted. For the
purposes of simplicity, the combined effects of laminate width and patch anchor length
on factor k4 is ignored. Such a simplification is expected to result in sightly
conservative predictions when patch anchor lengths of less than 300mm are used and
have no effect where anchor lengths of 300mm or greater are adopted in design.
The results due to variations of concrete strength, which were investigated in the finite
element models, were also predicted by the theoretical model to a good level of
accuracy. The model also shows a linear correlation between the concrete strength and
laminate strain achieved prior to debond in accordance with the finite element data.
When examining the effects of varying patch anchor width on anchorage performance,
the model provided predictions in failure load which were within 9% of the average
values for anchor type 4. Unfortunately, no experimental data was available for patch
anchor widths less than 220mm for evaluation.
A major constituent which distinguished the results for experimental stages 1 and 2 was
the laminate thickness used (2mm and 1.4mm). The formulations adopted in the
proposed model provided the necessary adjustments to the maximum shear strength of
the interface and the effective anchorage length resulting in a lower strain in the FRP
prior to failure where a higher laminate thickness was used. This was consistent with the
expected behaviour which resulted in reasonable predictions.
Cha
pter
9 –
Dev
elop
men
t of P
atch
Anc
hor P
redi
ctio
n M
odel
213
Tabl
e 9.
2– S
umm
ary
of e
xper
imen
tal a
nd n
umer
ical
pre
dict
ions
, ver
ified
with
the
prop
osed
anc
hora
ge st
reng
th m
ode
Spec
imen
Wid
th o
f Pa
tch
Anc
hor
(mm
)
Leng
th o
f Pa
tch
Anc
hor
(mm
)
Failu
re
Load
(k
N)
Max
FR
P st
rain
(
)
FRP
thic
knes
s (m
m)
FRP
Mod
ulus
(M
Pa)
FRP
Wid
th
(mm
)
Con
cret
e st
reng
th
(MPa
) r 1
r 2
r 3
r 4
u
(MPa
) L e
(mm
) P f
e
(kN
) fe
()
CO
V
(±%
)
()
Stag
e 1
WG
10
400
270
213
4900
2
1850
00
120
62
1.00
0 1.
00
1 0.
9 7.
81
186.
5 19
6.7
4431
-8
.3
WG
11
400
270
236.
9 53
00
2 18
5000
12
0 62
1.
000
1.00
1
0.9
7.81
18
6.5
196.
7 44
31
-20.
4 St
age
2
1.
1 40
0 30
0 13
1 44
06
1.4
2100
00
100
69.2
1.
050
0.83
1
1 7.
94
163.
6 14
2.1
4832
7.
8 1.
2 40
0 30
0 14
0.2
4922
1.
4 21
0000
10
0 69
.2
1.05
0 0.
83
1 1
7.94
16
3.6
142.
1 48
32
1.3
2.1
400
250
111
3819
1.
4 21
0000
10
0 69
.2
1.05
0 0.
83
1 0.
83
7.94
16
3.6
118.
4 40
27
6.2
2.2
400
250
128.
1 43
28
1.4
2100
00
100
69.2
1.
050
0.83
1
0.83
7.
94
163.
6 11
8.4
4027
-8
.2
3.1
300
300
151.
6 53
78
1.4
2100
00
100
69.2
1.
050
0.83
1
1 7.
94
163.
6 14
2.1
4832
-6
.7
3.2
300
300
138.
5 48
01
1.4
2100
00
100
69.2
1.
050
0.83
1
1 7.
94
163.
6 14
2.1
4832
2.
5 3.
3 30
0 30
0 15
8.8
5600
1.
4 21
0000
10
0 69
.2
1.05
0 0.
83
1 1
7.94
16
3.6
142.
1 48
32
-11.
8 3.
4 30
0 30
0 13
9.1
5091
1.
4 21
0000
10
0 69
.2
1.05
0 0.
83
1 1
7.94
16
3.6
142.
1 48
32
2.1
4.1
200
300
140.
6 49
50
1.4
2100
00
100
69.2
1.
050
0.83
0.
1 7.
94
163.
6 11
3.7
3866
-2
3.7
4.2
200
300
119.
9 45
04
1.4
2100
00
100
69.2
1.
050
0.83
0.
1 7.
94
163.
6 11
3.7
3866
-5
.5
4.3
200
300
112.
5 41
24
1.4
2100
00
100
69.2
1.
050
0.83
0.
1 7.
94
163.
6 11
3.7
3866
1.
0 4.
4 20
0 30
0 12
3.6
4514
1.
4 21
0000
10
0 69
.2
1.05
0 0.
83
0.1
7.94
16
3.6
113.
7 38
66
-8.8
St
age
1-FE
M P
aram
etri
c st
udy
W
G10
(32
Mpa
) 40
0 27
0 16
1.2
3632
2
1850
00
120
32
0.79
5 1.
00
1 0.
9 7.
81
186.
5 15
6.3
3521
-3
.1
WG
10 (4
5 M
pa)
400
270
187.
7 42
27
2 18
5000
12
0 45
0.
884
1.00
1
0.9
7.81
18
6.5
173.
8 39
15
-8.0
W
G10
(62
Mpa
) 40
0 27
0 21
3 47
97
2 18
5000
12
0 62
1.
000
1.00
1
0.9
7.81
18
6.5
196.
7 44
31
-8.3
St
age
2-FE
M P
aram
etri
c st
udy
Ty
pe 3
(32
Mpa
) 30
0 30
0 10
9.5
3725
1.
4 21
0000
10
0 32
0.
795
0.83
1
1 7.
94
163.
6 10
7.6
3659
-1
.8
Type
3 (4
5 M
pa)
300
300
125.
1 42
56
1.4
2100
00
100
45
0.88
4 0.
83
1 1
7.94
16
3.6
119.
6 40
69
-4.6
Ty
pe 3
(69.
2 M
pa)
300
300
150.
1 51
05
1.4
2100
00
100
69.2
1.
050
0.83
1
1 7.
94
163.
6 14
2.1
4832
-5
.7
Ave
rage
-5.2
Chapter 9 – Development of Patch Anchor Prediction Model
214
9.6 Summary
A theoretical strength prediction model has been developed for FRP patch anchored
joints, based on the results derived from experimental data and finite element parametric
studies. The model was capable of predicting patch anchor response, when varying
parameters such as: concrete strength, laminate width, laminate thickness, laminate
modulus, patch anchor length and patch anchor width. In addition, the model has been
verified to estimate the maximum laminate strains and loads reached prior to debond to
a reasonable level of accuracy.
Chapter 10 – Conclusion
215
10 CHAPTER 10 – CONCLUSION
The strengthening of existing reinforced concrete structures using fiber reinforced
polymers (FRP’s) as externally bonded reinforcement is gaining increasing attention
due to the materials superior mechanical properties and light weight. However, a serious
limitation in the use of FRP as a strengthening material comes from separation of the
FRP from the concrete surface by premature debonding at a strain level which is well
below the ultimate tensile strength of the material. Therefore, the focus of this
dissertation has been the research and development of new and efficient anchorage
systems to improve the strength utilization of FRP laminates bonded to concrete.
A state of the art review was presented which compiled the extensive amount of
experimental data on the various form of anchorages investigated over the past decade.
The data was consolidated and tabulated based on the anchorage type, material
properties, test configuration and maximum fiber elongation reached prior to debond.
The classification of data resulted in each type of anchorage being assigned an
anchorage effectiveness factor so that anchorage performance could be rated. For
flexural strengthening applications, it was found that the application of anchorages to
the ends of FRP laminate or sheet was effective in preventing the failure mechanism of
end debond. However, for the prevention of intermediate flexural and shear crack
induced debonding, anchorage throughout the span was also needed. Of the various
forms of anchorages examined, metallic anchorages were found to be the most effective
in preventing end debond, followed by U-jackets and FRP spike anchors. However for
shear strengthening applications metallic anchorages were found to be the least
effective.
Following the thorough review of the existing forms of anchorages available, it was
found that the majority were limited by a labour intensive installation process, subject to
corrosion and ongoing maintenance or required mechanical fasters. The primary
objective of the current research was to devise a new form of FRP anchorage which was
highly effective in the prevention of debonding, non-destructive, low maintenance and
easy to install. Anchorages in the forms of mechanical substrate strengthening and patch
anchors consisting of unidirectional and bidirectional fibers were conceived
Chapter 10 – Conclusion
216
conceptually and examined via a 2 stage experimental program, followed by extensive
numerical simulations and parametric studies.
The first stage of the experimental study consisted of improving the substrate properties
to which the FRP is bonded over the anchorage region by the introduction of a
mechanical chase cut into the concrete cover. The chase was effective in improving the
strength of the substrate, thereby shifting the debonding failure plane from between the
concrete-adhesive layer to the adhesive-FRP layer, which resulted in failure at a higher
load. The effect of the chase was a 95-100% increase in ultimate capacity, a 118%
increase in bond stress and 83-93% increase in the maximum strain level reached prior
to failure.
Although, the mechanical chase was effective in improving the anchorage strength, the
remainder of the experimental programme focused on non-destructive forms of anchors,
namely, unidirectional and bidirectional fiber patch anchors. Of the six types of
anchorages tested in stage 1 of the experimental programme, the use of bidirectional
fiber patch anchors was proven to be the most effective in increasing the anchorage
strength by up to 195%. Such a large increase in anchorage strength was achieved by
the patch anchors ability to distribute the adhesive-to-concrete bond stresses, typically
localised to the width of the FRP laminate, over a wider area of concrete.
Based on the results from the stage 1 study, a further experimental program was
designed in stage 2, with a specific focus on investigating the bidirectional fiber patch
anchors in more detail. A further 15 full scale anchorage specimens were tested with
varying parameters such as patch anchor sizes, laminate thickness and concrete strength,
resulting in sufficient experimental data for the basis of further finite element
simulations.
The finite element simulations, consisted of 3D nonlinear models, capable of cracking
and crushing response and replication of FRP debonding via the definition of interface
elements between the adhesive and the concrete which were calibrated to a predefined
bond-slip law derived from the experimental results. The models were successfully
calibrated with the experimental data and verified using all specimens tested resulting in
Chapter 10 – Conclusion
217
good predictions of the pre-peak and post-peak response of the joints. Furthermore,
sensitivity and parametric studies were performed to evaluate the influence of several
key parameters and the results were used to expand the experimental results to
encompass anchorage strength predictions for a wider range of concrete strengths. As a
result, an approximately linear relationship was discovered relating the strength of the
concrete and the maximum fiber elongation reached prior to debond for the patch
anchored joints. Finally, design formulations were proposed for patch anchor strength
predictions which were later verified with the experimental results.
It is recommended that future study should focus on the construction of large scale
shear strengthened RC beams, with FRP shear ligatures anchored using the patch
anchored developed herein.
References
218
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List of Publications
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LIST OF PUBLICATIONS
Journals:
Kalfat, R, Al-Mahaidi, R and Smith, S.T (2013). "Anchorage Devices used to improve the Performance of Reinforced Concrete Beams Retrofitted with FRP Composites: A-State-of-the-Art-Review." Journal of Composites for Construction 0(ja): 223.
Kalfat R and Al-Mahaidi R (2010). "Investigation into bond behaviour of a new CFRP anchorage system for concrete utilising a mechanically strengthened substrate." Journal Composite Structures 92(11): 2738-2746.
Al-Mahaidi, R and Kalfat R (2011). "Investigation into CFRP plate end anchorage utilising uni-directional fabric wrap." Journal of Composite Structures 93(2): 821-830.
Al-Mahaidi, R and Kalfat R (2011). "Investigation into CFRP laminate anchorage systems utilising bi-directional fabric wrap." Journal of Composite Structures 93(4): 1265-1274.
Kalfat R and Al-Mahaidi R (2013). “Numerical and Experimental Validation of FRP Patch Anchors used to improve the Performance of FRP Laminates Bonded to Concrete.” Journal of Composites for Construction, IIFC 10th Anniversary Issue, accepted for publication.
Conference papers and magazines:
Kalfat R and Al-Mahaidi R (2013). “Experimental and Numerical Investigation of Patch Anchors used to Enhance the Performance of FRP Laminates in Concrete Structures.” Article, Concrete in Australia Magazine, August 2013
Kalfat R and Al-Mahaidi R (2013). “Size Effect of Bi-directional Fibre Patch Anchors Used to Enhance the Performance of FRP Laminates.” FRPRCS-11, International Symposium on Fiber Reinforced Polymer Reinforcement for Reinforced Concrete Structures, Guimaraes , Portugal, June 2013 .
Kalfat R and Al-Mahaidi R (2012). “Finite Element Investigation of FRP Laminates Anchored using multi-layered Bi-directional Fibres.” The 6th International Conference on Advanced Composite Materials in Bridges and Structures, ACMBS-VI, Kingston, Ontario, Canada, from 22-25 May 2012.
Kalfat, R, Al-Mahaidi, R & Williams, G 2011, 'Investigation of efficient anchorage systems for shear and torsional retrofitting of box girder bridges', Proceedings, 10th International Symposium on Fiber Reinforced Polymer for Reinforced Concrete Structures (CD-Rom), FRPRCS-10, Tampa, Florida, USA, 2-4 April.
List of Publications
231
Kalfat R, Al-Mahaidi R and Williams G (2011). "The Application of FRP Anchorage systems in the Retrofitting of the Westgate bridge Project” Article, Concrete in Australia Magazine, Feb 2011
Kalfat R, (2008). "The Strengthening of Post-tensioned slabs using CFRP Composites at White City, London." Structural Faults and Repair, 12th international Congress and Exhibition, Edinburgh, 2008
Al-Mahaidi R, Kalfat R and Williams G (2011).”The use of innovative FRP Anchorages to improve the performance of Box Girder Bridge retrofit projects” First Middle East conference on Smart Modelling, Assessment and Rehabilitation of Civil Infrastructure. 8-11 February 2011, Dubai, UAE
Williams G, Al-Mahaidi R, Kalfat R (2011). “Carbon Fibre Retrofitting of the West Gate Bridge” Article, Concrete in Australia Magazine, Feb 2011
Williams G, Al-Mahaidi R and Kalfat R (2011). " The West Gate Bridge: Strengthening of a 20th Century Bridge for 21st Century Loading." Proceedings, 10th International Symposium on Fiber Reinforced Polymer for Reinforced Concrete Structures (CD-Rom), FRPRCS-10, Tampa, Florida, USA, 2-4 April.
Williams, G, Al-Mahaidi, R, Kalfat, R, (2011). "Strengthening of the West Gate Bridge, Melbourne, Australia." IIFC FRP International 8(3): 3-4.