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Investigation of Creep Rupture Phenomenon in Glass Fibre Reinforced Polymer
(GFRP) Stirrups
by
Kanwardeep Singh Johal
A thesis submitted in conformity with the requirements
for the degree of Master of Applied Science
Department of Civil Engineering
University of Toronto
© Copyright by Kanwardeep Singh Johal 2016
ii
Investigation of Creep Rupture Phenomenon in Glass Fibre Reinforced Polymer (GFRP)
Stirrups
Kanwardeep Singh Johal
Master of Applied Science
Department of Civil Engineering
University of Toronto
2016
Abstract
Glass Fibre-Reinforced Polymer (GFRP) bars offer a feasible alternative to typical steel
reinforcement in concrete structures where there are concerns of corrosion or magnetic
interference. In order to design safe structures for a service life of 50 to 100 years, the long-
term material properties of GFRP must be understood. Thirty GFRP stirrups of three types
were tested under sustained loading to investigate creep rupture and modulus degradation
behaviour. The time to failure under varying sustained loads was used to extrapolate the safe
design load for typical service lives. It was found that shear critical beams with shear
reinforcement designed in accordance with CSA-S806 and ACI-440 provisions may be at risk
of premature failure under sustained design loads. Analysis was based on finite element
modelling and previously tested beams. Additionally, no moduli degradation was observed in
this study. A cumulative weakening model was developed to potentially take into account
fatigue loading.
iii
Acknowledgements
This thesis would not have been possible without the contribution of many individuals at the
University of Toronto in the Department of Civil Engineering and external organizations.
Firstly, I would like to thank my supervisor, Professor Shamim A. Sheikh, not only did he give
me the opportunity to enter into the world of research but also the guidance and support to
navigate through it. Secondly, I would like to thank Professor Oh-Sung Kwon for taking the
time to read my thesis and providing feedback.
I would like to thank all of the technical staff in the University of Toronto Structural
Laboratory. Bryant Cook and Michel Fiss for showing me that anything can be taken apart and
put back together. Xiamong Sun, John MacDonald, and Giovanni Buzzeo for helping with
testing and machining. Renzo Basset for managing my project and listening to all my concerns.
Alan McClenaghan for being a friend and taking the time to share knowledge and stories.
My research experience would have been nothing without the friends and mentors I gathered
along the way. Thank you David Johnson for laying the foundation for this research project
and providing me with the answer key to Professor Sheikh’s questions. Thank you Arsalan
Tavassoli for keeping me sane, and sharing in the pain and laughter that research brings.
Additionally, I would like to show my gratitude towards Brandon Gemme, Guilherme Cestaro
and Weijie Liu for their hard work, muscle and attention to detail, which was very much
required in the structural lab.
I would also like to show my appreciation to the University of Toronto for not only providing
the resources to allow this project to materialize, but also for being my home for six years. I
would also like to thank NSERC and OGS for their financial support, which made this work
possible.
Finally, I would like to thank my family for all of their support and creating a culture of learning
and education at home. My parents, Kanwaljit and Manjit Johal, for putting my education
before all else. My older sister, Ramdeep Johal, for letting me borrow books with her library
pass and introducing me to Dr. Suess. My younger sister, Tarndeep Johal, for letting me watch
“National Geographic” on the television although “Dragontales” was airing. My grandfather,
Gulwant Toor, for teaching me Punjabi, properly. My grandmother, Jagjit Toor, for making
sure I was always well fed. My significant other, Manjot Gosal, for making me forget my
worries and showing me how to be carefree again.
iv
Dedicated in memory of my father, Manjit Johal, a true scholar.
v
Table of Contents
Table of Figures ...................................................................................................................... vii
Nomenclature ............................................................................................................................. x
1 Introduction ........................................................................................................................ 1
1.1 General ........................................................................................................................ 1
1.2 GFRP ........................................................................................................................... 2
1.2.1 Manufacturing ...................................................................................................... 3
1.2.2 Durability ............................................................................................................. 4
1.3 Objectives and Scope .................................................................................................. 6
1.3.1 Time to Failure ..................................................................................................... 7
1.3.2 Predication of Failure ........................................................................................... 7
1.3.3 Bend Strength....................................................................................................... 7
2 Literature Review............................................................................................................... 9
2.1 Relevant Work............................................................................................................. 9
2.1.1 Creep behaviour of pultruded GFRP elements – Part 1: Literature review and
experimental study ............................................................................................................. 9
2.1.2 Creep and durability of sand-coated glass FRP bars in concrete elements under
freeze/thaw cycling and sustained loads .......................................................................... 11
2.1.3 Creep rupture of a GFRP composite at elevated temperatures .......................... 13
2.1.4 Effect of Sustained Load and Environment on Long-Term Tensile Properties of
Glass Fibre-Reinforced Polymer Reinforcing Bars ......................................................... 14
2.1.5 Predicting the tensile strength and creep-rupture behaviour of pultruded glass-
reinforced polymer rods ................................................................................................... 16
2.1.6 Durability tests performed on straight ComBAR GFRP bars with standard
coating d=16mm Report No: 116 05 G ........................................................................... 18
2.1.7 Fibre reinforced polymer shear reinforcement for concrete members: behaviour
and design guidelines ....................................................................................................... 20
2.1.8 Bend Strength of FRP Stirrups: Comparison and Evaluation of Testing Methods
…………………………………………………………………………………22
2.1.9 Investigation of Glass Fibre Reinforced Polymer (GFRP) Bars as Internal
Reinforcement for Concrete Structures ........................................................................... 24
2.2 Relevant Codes, Clauses and Standards .................................................................... 28
2.3 Literature Review Summary ..................................................................................... 32
3 Experimental Program ..................................................................................................... 33
3.1 Material Properties .................................................................................................... 34
3.1.1 Concrete ............................................................................................................. 35
3.1.2 Reinforcing Steel ............................................................................................... 36
3.1.3 Glass Fibre Reinforcing Polymers Stirrups ....................................................... 37
vi
3.2 Construction of Specimens........................................................................................ 46
3.3 Instrumentation and Testing Equipment ................................................................... 47
3.3.1 Strain Gauges ..................................................................................................... 47
3.3.2 Load Cells .......................................................................................................... 50
3.3.3 Hydraulic Jacks .................................................................................................. 51
3.4 Test Observations ...................................................................................................... 53
4 Analysis of Data ............................................................................................................... 55
4.1 Stirrup 100% Tests .................................................................................................... 55
4.1.1 Time to failure (TTF) for monotonic loading .................................................... 60
4.2 Creep Rupture Results ............................................................................................... 62
4.2.1 S Series............................................................................................................... 64
4.2.2 Pt Series ............................................................................................................. 65
4.2.3 P Series............................................................................................................... 66
4.3 Creep Rupture and Sustained Loads ......................................................................... 67
4.3.1 Finite Element Modelling - VecTor 2 ................................................................ 74
4.4 Modulus Degradation and Creep Strain .................................................................... 82
4.4.1 Creep Strain ....................................................................................................... 83
4.4.2 Moduli Degredation ........................................................................................... 93
4.5 Cumulative Weakening Model................................................................................ 100
5 Conclusion ..................................................................................................................... 105
6 Recommendations for Future Work............................................................................... 107
7 References ...................................................................................................................... 108
Appendix A - Test Results ......................................................................................................... 1
P Series Strain and Applied Load versus Time ..................................................................... 1
Pt Series Strain and Applied Load versus Time .................................................................... 5
S Series Strain and Applied Load versus Time ................................................................... 10
S Series - Strain vs Stress..................................................................................................... 15
Pt Series - Strain vs Stress ................................................................................................... 21
Appendix B – Measurement Equipment Information .............................................................. 24
Appendix C - Strain Gauge - TML UFLA 2-350-11 ............................................................... 26
vii
Table of Figures
Figure 1-1 - Corrosion damage to underside of Gardiner Expressway (Beaty, 2012) .............. 2
Figure 1-2 - Manufacturing process for GFRP bars (Schöck, 2009) ......................................... 3
Figure 1-3 - Chemical composition of glass fiber (Left), single glass fibre (Right) viewed
through electron microscope (Ehrenstein, 2007) ....................................................................... 3
Figure 1-4 - Micro-cracking of resin matrix due to alkali attack (Benmokrane, 2002)............. 5
Figure 1-5 - Long-term reduction of tensile strength for glass fibres under sustained loading
(Benmokrane, 2002) .................................................................................................................. 6
Figure 2-1 - Theoretical creep strain domains for FRP materials (ASTM D2990-09) .............. 9
Figure 2-2 - Creep Deformation vs. Time for coupon tests (Sa et al. 2011) ........................... 10
Figure 2-3 - Creep Deformation vs. Time for beams (Sa et al. 2011) ..................................... 11
Figure 2-4 - Beam Schematic, Plan and Section (Laoubi, El-Salakawy, and Benmokrane,
2006) ........................................................................................................................................ 12
Figure 2-5 - Sustained Load deflection and strain for beams (Laoubi, El-Salakawy, and
Benmokrane, 2006) .................................................................................................................. 12
Figure 2-6 - Creep strain regression (Strain vs. Time) (Dutta and Hui, 2000) ........................ 13
Figure 2-7 - Sustained load test setup (Nkurunziza et al. 2005) .............................................. 14
Figure 2-8 - Creep Strain for bars exposed to de-ionized water (Nkurunziza et al. 2005) ...... 14
Figure 2-9 - Creep Strain for bars exposed to alkali solution (Nkurunziza et al. 2005) .......... 15
Figure 2-10 - Theorized domains of failure for FRP's under sustained loading (Nkurunziza et
al. 2005) ................................................................................................................................... 15
Figure 2-11 - Failure band for stochastic failure of specimens under sustained loading using
Weibull Distribution (Franke and Meyer, 1992) ..................................................................... 17
Figure 2-12 - Test setup for sustained load tests on straight GFRP bars (Weber, 2005) ......... 18
Figure 2-13 - Time to Failure curves for straight bars tested at various ambient temperatures
(Weber, 2005) .......................................................................................................................... 19
Figure 2-14 - Comparison of ComBAR sizes in creep rupture (Weber, 2005) ....................... 20
Figure 2-15 - Test setup for uniaxial loading of bent specimens (Sheata, Morphy, and
Rizkalla, 2000) ......................................................................................................................... 21
Figure 2-16 - Internal fibre configuration after bending (left), and stresses present on bend
region (right) (Ahmed et al. 2010) ........................................................................................... 22
Figure 2-17 - Standard B.5 test setup (Ahmed et al. 2010) ..................................................... 23
Figure 2-18 - Standard B.12 test setup (Ahmed et al. 2010) ................................................... 23
Figure 2-19 - Testing configuration (Plan) (Johnson and Sheikh, 2014) ................................ 24
Figure 2-20 - Beam Reinforcement Configuration (Section) (Johnson and Sheikh, 2014)..... 25
Figure 2-21 - Load - Deformation response for ComBAR stirrup reinforced beams (Johnson
and Sheikh, 2014) .................................................................................................................... 26
Figure 2-22 - Load - Deformation response for V-Rod stirrup reinforced beams (Johnson and
Sheikh, 2014) ........................................................................................................................... 26
Figure 2-23 - Strain readings at service loads (Johnson and Sheikh, 2014) ............................ 27
Figure 2-24 - Strain readings at failure (Johnson and Sheikh, 2014) ...................................... 28
Figure 2-25 - Template for TTF graph (adapted from ASTM D2990-09) .............................. 30
Figure 2-26 - Template for modulus degradation (adapted from ASTM D2990-09) .............. 30
Figure 2-27 - Time to failure example considering exposure conditions (Schöck, 2013) ....... 32
Figure 3-1 - Cylinder during testing (left), failed 100 mm diameter concrete cylinder (right)
.................................................................................................................................................. 35
Figure 3-2 - Concrete strength at 7, 28 and 62 days ................................................................ 36
Figure 3-3 - Steel reinforcement of specimens ........................................................................ 37
Figure 3-4 - Inside and outside diameters for sand coated and plastic coated bars ................. 38
viii
Figure 3-5 - Tensile coupon tests with steel couplers .............................................................. 38
Figure 3-6 - External MTS gauge ............................................................................................ 39
Figure 3-7 - Coupon after testing, failure in-between couplers ............................................... 39
Figure 3-8 - Deformed 12 mm coupon (S) test results ............................................................ 40
Figure 3-9 - Twisted 12 mm coupon (Pt) test results ............................................................... 40
Figure 3-10 - Sanded 12 mm coupon (P) test results ............................................................... 41
Figure 3-11 - GFRP Specimen dimensions as ordered from suppliers .................................... 42
Figure 3-12 - From left to right – P (left), Pt (centre), and S (right) Stirrups .......................... 43
Figure 3-13 - Bend diameter for P (Left) and S (Right) stirrups ............................................. 43
Figure 3-14 - 3 Point bend radius of Pt stirrup ....................................................................... 44
Figure 3-15 - Internal fibres of P stirrup bend region .............................................................. 45
Figure 3-16 - Internal fibres of Pt stirrup bend region ............................................................. 45
Figure 3-17 - Internal fibres of S stirrup bend region .............................................................. 46
Figure 3-18 - 3D View of formwork and specimen ................................................................. 47
Figure 3-19 - Strain gauge locations (typical) ......................................................................... 48
Figure 3-20 - Removal of sand coating and placement of 2mm gauge ................................... 49
Figure 3-21 - Gauge placement on bend, 4 and 5, and wrapping of foil tape .......................... 49
Figure 3-22 - Honeywell Model 53 Load Cell......................................................................... 50
Figure 3-23 - Formwork with specimen prior to cast .............................................................. 51
Figure 3-24 - Casting of all 30 specimens ............................................................................... 51
Figure 3-25 - 200 kN hydraulic cylinders ................................................................................ 52
Figure 3-26 - Test specimen with hydraulic cylinder, and potentiometers .............................. 52
Figure 3-27 - Pneumatic to hydraulic pump with pressure switch .......................................... 53
Figure 3-28 - Specimens under sustained load inside steel frame ........................................... 54
Figure 4-1 - Specimen P3 - Strain and Applied Load vs. Time ............................................... 55
Figure 4-2 - Specimen Pt9 - Strain and Applied Load vs. Time .............................................. 56
Figure 4-3 - Specimen S4 - Strain and Applied Load vs. Time ............................................... 56
Figure 4-4 - Typical failure of P Series specimens at bend location (P3 pictured) ................. 57
Figure 4-5 - Pt 9 specimen failure at bend location ................................................................. 57
Figure 4-6 - Pt 2 after failure with bend pull out ..................................................................... 58
Figure 4-7 - Typical failure of S Series specimens at bend location (S4 pictured) ................. 58
Figure 4-8 - No damage or failures at other fully bonded end (typical) .................................. 60
Figure 4-9 - Applied load vs. Time to Failure for S, P and Pt Series ...................................... 63
Figure 4-10 - S-Series - Applied Stress vs. Time to Failure .................................................... 64
Figure 4-11 - Pt Series, Applied Stress vs. Time to Failure .................................................... 65
Figure 4-12 - P Series, Applied Load vs. Time to Failure ....................................................... 66
Figure 4-13 - Stirrup strain values for beams tested by Johnson and Sheikh (2014) .............. 68
Figure 4-14 - Location of strain gauge for tests conducted by Johnson and Sheikh (2014) .... 69
Figure 4-15 - TA/TP illustration ................................................................................................ 70
Figure 4-16 - TA/TP values for all specimens .......................................................................... 71
Figure 4-17 - P Series Applied Load vs. TTF with 1, 2 and 3 standard deviation levels ........ 72
Figure 4-18- Pt Series Applied Load vs. TTF with 1, 2 and 3 standard deviation levels ........ 72
Figure 4-19 - S Series Applied Load vs. TTF with 1, 2 and 3 standard deviation levels ........ 73
Figure 4-20 - Shear stress distribution along rectangular cross section................................... 74
Figure 4-21 - Formworks beam model .................................................................................... 74
Figure 4-22 - Load vs. Midspan Deflection responses from VecTor 2 models and experiment
.................................................................................................................................................. 76
Figure 4-23 - Failure region of JSV-40-40B (Johnson and Sheikh, 2014) .............................. 76
Figure 4-24 - Augustus shear reinforcement stress at SLS (450 kN) ...................................... 77
Figure 4-25 - Shear Stress vs Height (152 kN Pre-Cracking) ................................................. 78
ix
Figure 4-26 - Shear Stress vs Height (450 kN SLS) ................................................................ 78
Figure 4-27 - Shear Stress vs Height (902 kN prior to failure) ............................................... 79
Figure 4-28 - Load vs Midspan Deflection, VecTor 2 models and experiment ...................... 80
Figure 4-29 – Bend rupture in stirrup (JSC-100-50) (Johnson and Sheikh, 2014) .................. 80
Figure 4-30 - Augustus shear reinforcement stress at 450 kN (SLS) ...................................... 81
Figure 4-31 - Shear stress along height (450 kN SLS) ............................................................ 81
Figure 4-32 - Shear Stress along Height (1002 kN Failure) .................................................... 82
Figure 4-33 - Creep strain calculation illustration ................................................................... 83
Figure 4-34 - Pt2 Strain and Applied Load vs. Time ............................................................... 84
Figure 4-35 - S7 Strain and Applied Load vs. Time ................................................................ 84
Figure 4-36 - Specimen S10 Strain and Applied Load vs. Time ............................................. 85
Figure 4-37 - Specimen S1 Strain and Applied Load vs. Time ............................................... 85
Figure 4-38 - Specimen S1 Gauge 6 and Applied Load vs Time ............................................ 86
Figure 4-39 - Pt 7 Strain and Applied Load vs. Time .............................................................. 86
Figure 4-40 - Pt 10 Strain and Applied Load vs. Time ............................................................ 87
Figure 4-41 - P5 Strain and Applied Load vs. Time ................................................................ 87
Figure 4-42 - S8 Strain and Applied Load vs. Time ................................................................ 88
Figure 4-43 - S9 Strain and Applied Load vs. Time ................................................................ 89
Figure 4-44 - Specimen S9 Gauge 4 and 6 vs. Time ............................................................... 89
Figure 4-45 - Specimen P8 Strain and Applied Load vs. Time ............................................... 90
Figure 4-46 - Total Creep Strain vs. Time to Failure .............................................................. 92
Figure 4-47 - Creep rupture domains ....................................................................................... 93
Figure 4-48 - Theoretical stress vs. strain behaviour ............................................................... 94
Figure 4-49 - Pt8 - Stress vs. Strain (Gauge 3) ........................................................................ 95
Figure 4-50 - Pt2 - Stress vs. Strain (Gauge 3) ........................................................................ 95
Figure 4-51 - P5 - Stress vs. Strain (Gauge 3) ......................................................................... 96
Figure 4-52 - P6 - Stress vs. Strain (Gauge 3) ......................................................................... 97
Figure 4-53 - P9 - Stress vs. Strain (Gauge 3) ......................................................................... 97
Figure 4-54 - S1 - Stress vs. Strain (Gauge 6) ......................................................................... 98
Figure 4-55 - S2 - Stress vs. Strain (Gauge 4) ......................................................................... 99
Figure 4-56 - S8 - Stress vs. Strain (Gauge 7) ......................................................................... 99
Figure 4-57 - Equivalent time illustration for S8 ................................................................... 100
Figure 4-58 - S Series Applied Stress vs. Time to Failure with cumulative weakening model
................................................................................................................................................ 102
Figure 4-59 - P Series Applied Stress vs. Time to Failure with cumulative weakening model
................................................................................................................................................ 102
Figure 4-60 - Pt Series Applied Stress vs. Time to Failure with cumulative weakening model
................................................................................................................................................ 103
Figure 4-61 - Cumulative weakening comparison to TA/TP .................................................. 103
x
Nomenclature
ACI – American Concrete Institute
AFRP – Aramid fibre reinforcing polymers.
ASTM – American Society for Testing and Materials
Characteristic tensile strength (German Standard) - The strength for a 95% confidence
interval or the mean strength minus 1.64 standard deviations
CE – Environmental factor used in ACI 440-12R
CFRP – Carbon fibre reinforcing polymers
CHBDC – Canadian Highway Bridge Design Code
CSA – Canadian Standards Association
CWM – Cumulative weakening model
Design tensile strength (German Standard) - Characteristic value divided by a partial safety
factor of 1.3
db – Bar diameter, typically in mm
ɛo – Initial strain, see equation 2-1
ɛcr – Creep strain
fbend – Strength of FRP bar at bend location
f’c – Concrete strength at 28 days, determined from 100 mm diameter cylinders
fstr = ffu – Strength of FRP bar at straight location, or strength of straight FRP bar
φFRP - FRP Material resistance factor, CSA S806 calls for a value of 0.75
GFRP – Glass fibre reinforcing polymers
JSCE – Japanese Society of Civil Engineering
lt – Tail length of stirrup at hooked end
ns – Sustained load vs. time to failure slope, similar to R10 value and temporal slope
P – Pultrall standard, type of GFRP stirrup manufactured by Pultrall Inc., marketed as VROD
Pt – Pultrall Twist, type of GFRP stirrup manufactured by Pultrall Inc., marketed as VROD
S – Schoeck, type of GFRP stirrup manufactured by Schoeck GmbH, marketed as ComBAR
σi – Applied tensile stress, typically in MPa
t – Time, adjusted time
to – Reference time value
xi
Δti – Time increment for numerical integration, typically taken as 1 second
T – Ambient temperature, see equation 2-1
Temporal slope – See ns
Tb – Baseline time to failure for specimen at a particular stress level, see equation 2-2
To – Reference temperature, see equation 2-1
TTF – Time to failure for a specimen under a constant stress
τmax – Maximum shear stress
rb – Bend radius
R10 – Difference of sustained load value between log decade of time
URS – Ultimate rupture stress
1
1 Introduction
1.1 General
Reinforced concrete is a composite material made of concrete and, typically, steel bars.
Concrete in itself is a composite of cement powder, aggregate, water and admixtures.
Formulations of concrete can vary for required needs, such as low density concrete for lower
weight, or chloride resistant to protect against salt environments where exposure to chlorides
is high (e.g. parking garages). Furthermore, concrete has excellent compressive strength,
however, it lacks tensile strength. Hence steel or other reinforcement is passively embedded
into the concrete during construction. This reinforcement allows the composite material,
reinforced concrete, to resist both tensile and compressive forces.
Many reinforced concrete structures are subject to harsh environments which can cause
deterioration and therefore the structure is unlikely to reach its design service life. In Canada,
40% of bridges and parking garages that are older than 40 years old are structurally or
functionally deficient (ISIS, 2001). Likewise, approximately 230,000 national bridges in the
U.S. are corrosion ridden and considered structurally deficient (DCS, 2004). This corrosion
has led to repair costs exceeding $6 billion per year in Canada for parking garages alone while
in the U.S. the estimates reach $75 billion in repair costs each year (Benmokrane, 2005).
Deterioration of reinforced concrete structures will continue to grow and costs mount, therefore
preventative design measures are required for new structures so they do not share the same fate
as current structures.
Figure 1-1 shows the Gardiner Expressway exhibiting high levels of corrosion based spalling
due to the application of salts and de-icing chemicals. The expressway was completed in 1966;
at the time salt was not used for de-icing and therefore the Gardiner, like many other structures
from that era, was not designed with adequate corrosion protection.
2
Figure 1-1 - Corrosion damage to underside of Gardiner Expressway (Beaty, 2012)
1.2 GFRP
In efforts to develop corrosion-free concrete structures, GFRP bars have been considered as an
alternative to typical steel reinforcement. Although currently more expensive, GFRP has a
much higher strength to weight ratio than that of traditional steel reinforcement. Table 1-1
compares the nominal weight per meter of rebar for steel and GFRP (Schöck, 2009). The
average tensile strength of GFRP and steel bars were assumed 1200 and 400 MPa in calculating
strength to weight ratios in Table 1-1, column 4.
Table 1-1 Nominal weight comparison for GFRP and Steel (Schöck, 2009)
Bar diameter
(mm)
GFRP (kg/m) Steel (kg/m) (σ/W)GFRP/
(σ/W)Steel
8 0.14 0.40 8.6
12 0.30 0.89 8.9
16 0.52 1.58 9.1
25 1.22 3.85 9.5
32 1.95 6.31 9.7
However, questions have been raised about their chemical and mechanical durability; this will
be discussed in further detail in the following chapters.
3
1.2.1 Manufacturing
GFRP bars are composed of thousands of small glass fibres with a diameter of about 15 µm
and a resin, typically vinyl ester based. They are produced in what is known as a pultrusion
process, where fibers are pulled off spindles into a preformed die where it is infused with low
viscosity liquid resin. Once the resin has cured, the hardened bars are cut to size. Figure 1-2
illustrates the typical pultrusion process for GFRP bars (Schöck, 2009). The individual fibres
are created by melting silica sand, limestone, and other minerals in large furnaces and extruding
the liquid through bushings (Bank, 2006). Once the fibres reach the required length they are
cut and bundled for use in the pultrusion process. Figure 1-3 displays the chemical composition
of glass fiber (left) and a single fiber as viewed through an electron microscope (right).
Figure 1-2 - Manufacturing process for GFRP bars (Schöck, 2009)
Figure 1-3 - Chemical composition of glass fiber (Left), single glass fibre (Right) viewed
through electron microscope (Ehrenstein, 2007)
4
GFRP can be manufactured from different types of glass; currently manufacturers typically use
ECR Glass because of its combination alkali resistance and lower cost. The various types of
glass fibres and their properties have been summarized in Table 1-2.
Table 1-2 - Commonly used glass fibre types (Johnson, 2009)
Polymers are formed through a process called polymerization, in which monomers react with
one another to form linear chains, branches or cross-linked three-dimensional networks.
Thermosetting polymers are preferred for current GFRP manufacturing, as they form a rigid
impervious three-dimensional cross-linked matrix (Moukwa, 1996). This is ideal for their
purpose as it provides solvent resistance, and high strength. Cross-linking is the establishment
of links between different chains and branches; this causes a rapid increase in viscosity and
hardness. The point at which this occurs is called the gel point; the resin is irreversibly
transformed from a liquid into a gel (Moukwa, 1996). Additionally, manufacturers can add
reactive monomers, typically styrene based, to increase crosslinking as it may not be possible
through vinyl esters alone because full curing may take too long. As crosslink density of
thermosets increase, they can become brittle and fracture at lower strains.
Various forms of thermosetting resins can be used in this application, such as polyester, epoxy,
and vinyl ester resins. Polyester resins are easier to cure, however, lack chemical and
mechanical durability. Epoxy based resins are seen as the most durable, but require high
temperatures to cure. Vinyl-ester resins combine some of the better properties of polyester and
epoxy resins, and are currently used in GFRP bars (Johnson, 2009).
1.2.2 Durability
Chemical
Although glass fibers are resistant to chloride attack, as they are non-metallic, they are not
resistant to alkaline ions. The degradation of the glass fibers occurs from the solution, which
forms within concrete air voids that can reach a pH of 13.5 (Benmokrane, 2005). Equation 1-
5
1 shows that free hydroxide (OH-) can degrade the silica based glass fibers (SiO2), causing the
formation of SiO solution and SiOH in gel form (Yilmaz, 1991).
SiO2 + OH- -> 2 SiOH (solid) + SiO (solution) [1-1]
GFRP bars obtain their overall tensile strength from the glass fibers, the resin only transfers
load from one fiber to another. The mechanism for the attack is only possible if the hydroxide
solution can permeate through the resin matrix and reach the glass fibers. If this resin is
impermeable it can provide adequate long-term protection for the fibers, allowing the fibers to
adequately resist the mechanical loads.
The resin matrix is also susceptible to chemical attack from chlorides and hydroxides. The
vinyl ester resin is composed of long organic molecular chains formed during polymerization.
If this matrix has prolonged exposure to OH- and Cl- the bonds between carbon and oxygen
can be destroyed, leading to micro cracking (Benmokrane, 2005), as shown in Figure 1-4. A
damaged resin matrix causes a reduction in mechanical strength, and is more susceptible to
chemical ingress, accelerating fiber degradation.
Figure 1-4 - Micro-cracking of resin matrix due to alkali attack (Benmokrane, 2002)
Mechanical
Mechanical failure of GFRP bars can occur by any one, or a combination of the following
mechanisms: rupturing of fibers, cracking of the resin matrix, or loss of bond at the fiber-resin
interface. Early research has shown that creep deformation of GFRP is about twice that of steel.
Under an accelerated aging equivalent to 50 years of sustained loading, ultimate tensile strength
drops to 45% of no aging (Greenwood, 2002). However, this creep testing was done on
6
polyester resins, which is less durable than vinyl ester resins. Greater detail on mechanical
failure will be provided in the literature review section. Furthermore, it has been shown that
accelerated aging does not represent real world conditions and may be overly conservative.
Figure 1-5 shows the reduction in tensile strength of glass fibers conditioned to accelerated
aging (Benmokrane 2002).
Figure 1-5 - Long-term reduction of tensile strength for glass fibres under sustained loading
(Benmokrane, 2002)
1.3 Objectives and Scope
The main objective of the proposed research is to develop an understanding of bent GFRP bars
as shear reinforcement under sustained stresses. With the newer generation bent and straight
bars exhibiting adequate short term strength, the long-term performance remains a critical
concern. Considering the catastrophic nature of shear failure and the intended service life of
GFRP reinforced structures, the long-term performance of the shear reinforcement requires
thorough investigation. A clear understanding of the behaviour will assist in developing safe
design strengths and provide further information for discussions regarding a material resistance
factor.
Testing of 30 specimens (stirrups), three different types, from two different manufacturers was
conducted in order to determine the creep rupture response for GFRP reinforcing bars.
Specifically, GFRP bars bent into a stirrup shape used for shear reinforcement in beams.
Schöck GmbH provided 10 specimens. Pultrall Inc. provided the remaining 20 specimens, two
types of 10 bars each. Both manufacturers provided accompanying straight bars for coupon
testing.
7
1.3.1 Time to Failure
In order to quantify the behaviour of creep rupture, a “Time to Failure” relationship was
formed. Specimens were loaded monotonically in tension to failure, establishing the ultimate
load. This was followed by the sustained loading of the remaining specimens to percentiles of
the ultimate load, established previously. Intuitively, specimens with lower sustained loads
took longer to fail than the more heavily loaded ones. The times to failure for the specimens
covered 5 time epochs (1, 10, 100, and 1000 hrs). This relationship was plotted in order to
extrapolate the sustained load level that would cause failure at the 1 million hour mark
(nominally 100 years). The load level at 100 years minus three standard deviations, or an
alternative safety factor, would provide engineers safe sustained load design values.
1.3.2 Predication of Failure
This project aimed to go beyond simple charts that would provide an estimate for expected
failure times. We felt it was important to understand the mechanism by which creep rupture
was occurring and attempt to quantify and model it. This was accomplished by placing eight
strain gauges on each of the thirty specimens and measure strain values over time. Using
existing theories and knowledge about creep in polymeric compounds, it was possible to create
“Time to Failure” models for sustained loading. Additionally, if structures in the future were
to be ‘intelligent’ and be gauged over their lifetime, it would be possible to monitor structures
and be aware of any impending creep rupture induced failures.
1.3.3 Bend Strength
GFRP reinforcing bars are a composite material, one part glass fibre strands, and second part
an epoxy/resin. This material relies on the glass fibres to carry the load and resin to transfer
load from one strand to another. This system is effective when the strands remain co linear,
however, in bent sections deformities and voids cause a significant reduction in strength.
Although Vint and Sheikh (2015) have looked into bend strength and bonding, this project will
be specific to stirrups. The goal is to determine: percentage of the straight bar strength that
remains after bending them, factors causing a reduction in strength, and ultimately suggestions
for improvements in the manufacturing of bent bars.
To summarize, this project is designed to better understand GFRP stirrups as used for shear
and confinement reinforcement. This research project has tested the most current GFRP
8
products from two manufacturers, who have the largest market share, in an attempt to be as
relevant as possible. However, advancements in GFRP technology are being made quickly.
The quality of bars produced has improved significantly, for example, the modulus of GFRP
bars has gone from 40 GPa to 70 GPa over the last decade. So it is inevitable that this research
will age rather quickly, but we hope the methods used to analyse the state of the art materials
and the foundation built will stay relevant for many years to come.
9
2 Literature Review
2.1 Relevant Work
2.1.1 Creep behaviour of pultruded GFRP elements – Part 1: Literature review and
experimental study
Sa, Gomes, Correria, and Silvestre (2011)
The experimental program called for the testing of, small scale laminate GFRP beams (8 x 8 x
160 mm long) loaded to 20% to 80% of their Ultimate Rupture Strength (URS) and extruded
GFRP I beams (150 x 75 x 8 x 1800 mm long) loaded to 33% of their URS. The load was
maintained for the small scale laminate beam tests until failure occurred, while, for the extruded
beams the specimens were placed under a sustained load for a pre-determined duration and
then loaded monotonically until failure. Testing on the laminate GFRP beams showed that
heavily loaded specimens (60% to 80%) crossed the three stages of creep (primary, secondary
and tertiary as shown in Figure 2-1). All three stages of creep occur while the applied load is
constant. Primary creep is described as increase of total strain following a logarithmic function.
Secondary creep is described as the increase of total strain following a linear function. Tertiary
creep is described as increase of total strain following an exponential function. However, the
small specimens loaded to 50% and less of URS did not show significant creep strains (3% of
total strain). Results of creep tests showed higher creep deformation during the first 2000 hours,
after which the total strain remained nearly constant.
Figure 2-1 - Theoretical creep strain domains for FRP materials (ASTM D2990-09)
It was observed that the creep responses of the large beams (Figure 2-2) were similar to that of
the small-scale tests, and consistent mechanical behaviour between the two was achieved.
10
Figure 2-2 - Creep Deformation vs. Time for coupon tests (Sa et al. 2011)
From these tests it appears that GFRP specimens loaded at and above 60% of URS fall within
the domain of fibre rupture induced failure. Those specimens loaded between 50% and 60% of
URS underwent a matrix cracking failure. Specimens below 50% were in the diffusion failure
domain hence did not show significant creep throughout its life but showed minor tertiary creep
prior to failure (see Figure 2-3). It should be noted that this work deals with extruded GFRP
beams, which theoretically, would see inter-laminar shear forces and may not truly represent
the loading state of straight GFRP bars in tension. However, a good agreement was seen
between the extruded beams and coupon tests, and minimal shear deformation observed.
Therefore, these findings are applicable to the research undertaken by this thesis.
11
Figure 2-3 - Creep Deformation vs. Time for beams (Sa et al. 2011)
2.1.2 Creep and durability of sand-coated glass FRP bars in concrete elements under
freeze/thaw cycling and sustained loads
Laoubi, El-Salakawy and Benmokrane (2006)
Twenty-one concrete beams (1800 x 130 x 180 mm) reinforced with FRP composite bars were
tested in order to determine the combined and individual effects of freeze/thaw and sustained
loading (Figure 2-4). Specimens were exposed to 100, 200 or 360 freeze/thaw cycles (-20 °C
to +20 °C) either in an unstressed condition or loaded in four point bending. The sustained
tensile stress was set corresponding to 27% of the ultimate rupture strength of the GFRP bars
for a set time, at which point the beam was loaded until failure.
12
Figure 2-4 - Beam Schematic, Plan and Section (Laoubi, El-Salakawy, and Benmokrane,
2006)
The sustained load caused an initial GFRP strain (elastic) of 4900 micro-strain, more than
double the ACI 440.1R-06 creep strain limit of 2100 micro-strain. All beams, regardless of
temperature exposure, saw increased mid span deflection with time. The creep during the first
28 days was greater than the remaining 332 days (Figure 2-5).
Figure 2-5 - Sustained Load deflection and strain for beams (Laoubi, El-Salakawy, and
Benmokrane, 2006)
After 180 days, the mid-span deflection increased by an additional 10.5% to the initial elastic
deflection at the onset of loading. However, the long term creep strain measured in the
longitudinal GFRP bars was reported to be only an additional 1.9%. The larger mid deflection
is likely caused by the concrete undergoing creep deformation and not the GFRP
reinforcement.
The authors concluded that the ACI 440.1R-06 equation to predict long term deflections due
to creep and shrinkage over estimated creep strains by a factor of 5 and the stress limit of 0.20fu
is conservative.
13
2.1.3 Creep rupture of a GFRP composite at elevated temperatures
Dutta and Hui (2000)
A series of coupon tests on glass/polyester FRPs were performed on 6.35 mm thick specimens
following the ASTM D-3410 standard. The specimens were placed under sustained loads at
various temperatures in order quantify creep strain, modulus loss, poisons ratio, and ultimate
stress. We shall concern ourselves only with those tests conducted at room temperature (25 °C)
as it is of the most relevance to this study. Loading ranged from 60% to 80% of URS and TTF
ranged from 0 to 30 minutes. The researchers observed semi brittle failure with modulus
degradation at higher loads. Specimens underwent primary and secondary creep as seen in
Figure 2-6, however potential tertiary creep was not witnessed as tests lasting longer than 30
minutes were terminated. Again primary creep and secondary creep is defined as the increase
in total strain following a logarithmic and linear function, respectively.
Figure 2-6 - Creep strain regression (Strain vs. Time) (Dutta and Hui, 2000)
The authors used Findley’s equation (eqn. 2-1), modified with two empirically derived
constants p and β, to predict creep strains. A strong agreement was reached between the
experimental data and theoretical equation, albeit over a limited time span.
𝜀(𝑡) = [𝜀0 + 𝑝 (𝑡
𝑡0)𝛽(𝑇 𝑇0
⁄ )] [2-1]
14
2.1.4 Effect of Sustained Load and Environment on Long-Term Tensile Properties of Glass
Fibre-Reinforced Polymer Reinforcing Bars
Nkurunziza, Benmokrane, Debaiky, and Masmoudi (2005)
Twenty GFRP bars (E-Glass in vinylester matrix) 9.5 mm in diameter were tested over 417
days (10,000 hours) under a combination of a sustained load (25% or 38% of fu) and exposure
to a liquid (alkaline solution with a pH of 12.8 or de-ionized water pH 7.0) as shown in Figure
2-7. Once the exposure and sustained load time had elapsed, the bars were loaded until failure
to discover their residual strength and modulus. The residual strength of the bars exposed to
de-ionized water saw no significant reduction from the original values. Furthermore, the bars
exposed to alkaline saw a reduction of approximately 12%. The residual modulus ranged from
38.5 to 42.9 GPa, nearly the equivalent to the original value (Figure 2-8 and 2-9).
Figure 2-7 - Sustained load test setup (Nkurunziza et al. 2005)
Figure 2-8 - Creep Strain for bars exposed to de-ionized water (Nkurunziza et al. 2005)
15
Figure 2-9 - Creep Strain for bars exposed to alkali solution (Nkurunziza et al. 2005)
The reasoning for the low reduction in ultimate strength and modulus is due to the failure
domain associated with the applied sustained stress. FRP failures are categorized into three
transitional types, fibre rupture, matrix cracking, and diffusion (Figure 2-10). Fibre rupture
being when individual fibres either fail individually or delaminate from the resin and are no
longer able to transfer load. At moderate stress levels, the matrix suffers micro cracking. These
micro cracks eventually coalesce and when the resin matrix is unable to transfer load from one
fibre to another, a failure surface forms. Additionally, these micro cracks allow the permeation
of alkali solution into the bar, reducing its service life. At the diffusion stage, small voids can
form inside the bar but do not form cracks. Thus, there is no direct attack of the fibres from
surrounding media, failure can only occur by slow diffusion of alkali into the bar and failure at
extended periods of time. An applied stress of 25% to 38% falls within the region of diffusion,
and no damage to the matrix likely occurred, let alone the glass fibres.
Figure 2-10 - Theorized domains of failure for FRP's under sustained loading (Nkurunziza et
al. 2005)
16
The authors concluded that creep strain decays with time, and nearly all creep strain is observed
in the first 10,000 hours. Additionally, at 75 years of simulated aging under sustained stress,
GFRP bars experienced only an 8% strain increase (5% occurring in the first 10,000 h) in
addition to the initial strain, at the 38% stress level. Finally, the modulus of elasticity of the
bars was very stable and almost unaffected by the sustained stress levels used in this study
(25% to 38%). This is critical for the design of concrete elements reinforced with FRP bars
because the modulus is directly correlated to crack widths, deflection and other serviceability
concerns.
2.1.5 Predicting the tensile strength and creep-rupture behaviour of pultruded glass-
reinforced polymer rods
Franke and Meyer (1992)
The “Democratic Load Sharing Model”, with respect to FRPs, states that the sum of the load
carried by all individual fibres in a cross section is equal to the total load carried by said cross
section. Furthermore, the load carried by a fibre is determined by the local average stress, and
the stress across the total cross section may vary for several reasons, such as uneven loading
conditions, uneven fibre density or damaged fibres. Therefore, the loading on an individual
fibre may be higher than the average (e.g. the fibre adjacent is damaged and not carrying its
“fair-share”). This leads us to “The Cumulative Weakening Model” (CWM), which states that
a fibre failure will occur when it cannot handle its local stress condition, and upon failure, the
load will be redistributed to the remaining fibres. After an initial failure, the probability of a
local fibre being over stressed is increased, leading to more failures. This continues to occur
until all fibres are overstressed and a complete failure of the section occurs.
The described probabilistic, progressive failure typically follows a Weibull distribution. Franke
and Meyer (1992) applied this model to data produced by tensile tests on FRP bundled rods.
With the application of a probabilistic model, they were able to provide a range for TTF for a
given stress, allowing engineers to calculate a guaranteed tensile strength (mean strength minus
3 standard deviations) or an alternative safety margin (Figure 2-11).
17
Figure 2-11 - Failure band for stochastic failure of specimens under sustained loading using
Weibull Distribution (Franke and Meyer, 1992)
In order to account for variable loading conditions, equation 2-2 from the CWM was used to
calculate the total creep rupture stress that the specimen had incurred.
∑∆𝑡𝑖
𝑇𝑏(𝜎𝑖)
𝑗𝑖=1 = 1 [2-2]
Tb(σi) is the time which leads to failure for a particular sustained load, Δti, is the time for which
the particular sustained load was applied. When the sum of the various creep stresses equals 1,
creep rupture will occur. One may also convert a sustained load from one level to another using
equation 2-3.
𝑡1
𝑡2= (
𝜎2
𝜎1)𝑛𝑠
[2-3]
Where ns is the slope of the TTF curve as shown in Figure 2-11, and t1, t2, σ1 and σ2 are the
times and sustained stresses at each associated load, respectively. For more complex loading,
it is also possible to simply integrate stress with respect to time and determine an equivalent
time for a given stress level (equation 2-4).
𝑡1 =1
𝜎1𝑛𝑠 ∫ 𝜎(𝑡)𝑛𝑠𝑑𝑡
𝑡𝑏0
[2-4]
18
2.1.6 Durability tests performed on straight ComBAR GFRP bars with standard coating
d=16mm Report No: 116 05 G
Weber (2005)
Schöck GmbH conducted creep rupture tests on GFRP straight bars, marketed as ComBAR, of
varying diameters and temperatures. Bars were cast in concrete blocks and pushed apart by
hydraulic actuators in order to maintain a constant tensile load (see Figure 2-12). The load was
held until failure occurred. The failure times spanned 3 time epochs (10, 100, and 1000 hours).
The prescribed test method from CSA S806 Annex H calls for direct tension tests without the
presence of alkali, however, the authors decided to incorporate the effects of alkali exposure
by casting a concrete prism around the failure region.
Weber (2005) completed the tests on three bar sizes, 8, 16 and 25 mm at three temperatures,
30, 40 and 60 ºC. The data for the 16 mm straight bars at the various temperatures is plotted in
Figure 2-13, along with extrapolation to the 1 millionth hour (nominally 100 years).
Information on the linear regression is summarized in Table 2-1. CSA S806-12 requires for an
R2 value of a least 0.98 for the test to be considered valid. From the equation, the sustained
load that would cause failure in 100 years is equal to 602 MPa, approximately 46.1 % of the
ultimate tensile strength.
Figure 2-12 - Test setup for sustained load tests on straight GFRP bars (Weber, 2005)
19
Figure 2-13 - Time to Failure curves for straight bars tested at various ambient temperatures
(Weber, 2005)
Table 2-1 - Linear Regression of Creep Rupture Data for 16 mm diameter bars
It is also possible to determine the effect size may play on creep rupture strength by comparing
tests from the different bar sizes. Figure 2-14 provides a comparison of the three different bars,
with an ambient temperature of 60 ºC, and it can be seen that there is a strong agreement
between the data sets. It can, therefore, be concluded that the diameter of the bar plays little to
no role on creep rupture strength within this range of bar diameters.
20
Figure 2-14 - Comparison of ComBAR sizes in creep rupture (Weber, 2005)
2.1.7 Fibre reinforced polymer shear reinforcement for concrete members: behaviour and
design guidelines
Sheata, Morphy and Rizkalla (2000)
In this experimental program, the researchers tested FRP stirrups under uni-axial tension in
order to determine the bend strength as a function of bend radius and straight bar strength.
GFRP stirrups were cast into 50 MPa concrete (28 days) in two configurations (Type A and
Type B). The Type A setup was designed to measure the strength of hooked type stirrup, while
Type B was designed to measure the strength of a continuous stirrup (see Figure 2-15).
21
Figure 2-15 - Test setup for uniaxial loading of bent specimens (Sheata, Morphy, and
Rizkalla, 2000)
The 12 mm diameter GFRP stirrups possessed strength of 640 MPa, a modulus of 41 GPa, and
bend radius to diameter ratio of 4. According to the Japanese Society of Civil Engineering
(JSCE) equation 2-5, a bend radius to diameter ratio of 4 would produce a bend strength of
50%.
𝑓𝑏𝑒𝑛𝑑
𝑓𝑠𝑡𝑟= 0.05
𝑟𝑏
𝑑𝑏+ 0.30 [2-5]
Test results showed the bend strength equalled 49% (346 MPa) of the straight portion strength.
The authors called for a design limit of 0.4fy to be used for FRP stirrups, and a bend radius of
no less than 50 mm or a bend radius to diameter ratio of 4, whichever is greater.
22
2.1.8 Bend Strength of FRP Stirrups: Comparison and Evaluation of Testing Methods
Ahmed, El-Sayad, El-Salakawy, and Benmokrane (2010)
A comparison and evaluation of the current test methods used to determine the strength of FRP
bent bars was conducted. Test methods ACI 440.3R-04 B.12 and B.5 restrain the bend section
while applying a tensile load to the straight section in order to determine the capacity of the
bend. The bend strength is of concern because of its reduced strength due to the non-linearity
of fibres and the anisotropic characteristic of FRP. When the FRP bar is bent, the innermost
fibres kink and buckle to a higher degree than the outermost fibres, as shown in Figure 2-16
(left). This creates an area where there is a higher degree of voids and imperfections.
Furthermore, the FRP’s strength is highest in the parallel direction and weakest in the
transverse direction. Therefore, the shear force present at the end of the bend, Figure 2-16
(right), further reduces the ultimate tensile strength the stirrup can resist.
Figure 2-16 - Internal fibre configuration after bending (left), and stresses present on bend
region (right) (Ahmed et al. 2010)
Test method B.5, Figure 2-17, calls for the casting of a stirrup inside two concrete blocks, one
on each end. A hydraulic jack is placed in between the two blocks and loaded, generating a
tensile force. Test method B.12, Figure 2-18, works by placing a U-shaped FRP specimen
inside the rig, with two anchor points affixed to both terminal ends. A tensile load is applied to
the top frame causing a reaction against the anchors.
23
Figure 2-17 - Standard B.5 test setup (Ahmed et al. 2010)
Figure 2-18 - Standard B.12 test setup (Ahmed et al. 2010)
Twenty seven tests were conducted on GFRP specimens, 12 using method B.5 and 15 using
method B.12. The bars had diameters of 9.5 , 15.9, and 19.1 mm; the ratio of bend radius to
bar diamter was fixed to 4.0 for all tests.
24
Table 2-2 - Bend strength to straight bar strength comparison for test setups B.5 and B.12
fbend/fstraight
Diamter (mm) B.5 B.12
9.5 0.58 0.35
15.9 0.68 0.40
19.1 0.58 0.41
Table 2-2 summarizes the results of the experimental tests, and shows test method B.12
displays lower strength for the GFRP bends. The authors concluded that B.5 was a superior
test in that it yielded more accurate results when compared to JSCE’s prediction Furthermore,
test method B.5 is more representative of the actual loading and environmental conditions that
GFRP stirrups would see. It was concluded B.12 should only be used for quality control
purposes once a benchmark has been established. CSA S806 Annex D requires the testing of
GFRP specimens using test method B.5 as carried out in this project.
2.1.9 Investigation of Glass Fibre Reinforced Polymer (GFRP) Bars as Internal
Reinforcement for Concrete Structures
Johnson and Sheikh (2014)
The experimental work included the testing of 24 large scale beams (400 x 650 x 3650 mm) in
order to determine the behaviour of GFRP transverse reinforcement. The beams were designed
to be shear critical and were placed under three point loading (Figure 2-19). Of these 24 tests,
we shall concern ourselves with only eight of the beams; those of which are reinforced with
GFRP stirrups cast in regular and high strength concrete. Table 2-3 summarizes the material
properties for the GFRP bars used to reinforce the beams tested.
Figure 2-19 - Testing configuration (Plan) (Johnson and Sheikh, 2014)
25
Table 2-3 - Material properties of select beams tested in study (Johnson and Sheikh, 2014)
The six beams were reinforced with GFRP stirrups, with three different reinforcement ratios
(0.22%, 0.40% and 0.50%) from two different manufacturers (C and B1). Figure 2-20 shows
the cross sections for the different beams and their reinforcement detailing.
Figure 2-20 - Beam Reinforcement Configuration (Section) (Johnson and Sheikh, 2014)
Load deflection curves are shown in Figure 2-21 and Figure 2-22. Although an increase in
shear reinforcement increased the ultimate capacity of the beams, the initial deflection was
similar (the flexural reinforcement was constant). At the SLS deflection limit of 10 mm
(Assuming L./360) the corresponding load was approximately 400 kN (450 kN was considered
to be the serviceability load level).
26
Figure 2-21 - Load - Deformation response for ComBAR stirrup reinforced beams (Johnson
and Sheikh, 2014)
Figure 2-22 - Load - Deformation response for V-Rod stirrup reinforced beams (Johnson and
Sheikh, 2014)
27
Stirrup strain values, average and peak, have been provided for the 450 kN load level in Figure
2-23. This is of significant importantance as it provides insight into the sustained stress levels
GFRP stirrups would be subjected to. Only the minimally reinforced beams had stirrups which
exceeded the ACI limit of 0.004 and none exceeded the CSA S806 limit of 0.005. Figure 2-24
shows the strain values at bend and straight locations at failure.
Figure 2-23 - Strain readings at service loads (Johnson and Sheikh, 2014)
28
Figure 2-24 - Strain readings at failure (Johnson and Sheikh, 2014)
2.2 Relevant Codes, Clauses and Standards
The design and qualification of FRP materials as reinforcement for concrete is governed by
CSA S806-12 (Design), and CSA S807-10 (Qualification) in Canada. Additionally, the
Canadian Highway Bridge Design Code, S6-14, has its own provisions for the design of bridges
with FRP. S6 borrows heavily from the ACI 440.1R-15 and previous editions, which sets the
standards for FRP design for the U.S. In this section relevant codes and clauses have been
provided along with technical information from Schöck GmbH.
CSA S807-10
Table 3 - Minimum strength at the bend shall be at least 45% of the minimum strength of
straight FRP bars manufactured by the same process as the bent bars
CSA S806-12
7.1.6.3 - For non-prestressed reinforcement, the resistance factor, φFRP, shall be taken as φFRP
= 0.75.
29
7.1.2.2 - The maximum stress in FRP bars or grids under loads at serviceability limit state
shall not exceed the following fraction of the characteristic tensile strength:
(a) for AFRP, 0.35;
(b) for CFRP, 0.65; and
(c) for GFRP, 0.25.
7.1.2.3 - The maximum strain in GFRP tension reinforcement under sustained service loads
shall not exceed 0.002.
9.5 - The tail length of a bent bar , lt, should not be less than 12db . The bend radius, rb ,
should not be less than 3db.
CSA S6.2012
Table 16.2 - φFRP for pultruded FRP shall be taken as 0.55
16.8.3 - Non-prestressed reinforcement
The maximum stress in FRP bars or grids under loads at SLS shall not be more than FSLS x
fFRPu, where FSLS is as follows:
(a) for AFRP: 0.35;
(b) for CFRP: 0.65; and
(c) for GFRP: 0.25.
ACI 440-12R
To avoid creep rupture of FRP reinforcement under sustained loading (servicability), ACI 440,
places a stress limit of 0.20ffu. For ultimate limit state design, the strain on FRP shear
reinforcement should not exceed 0.004, in order to preserve agregate interlock for concrete
shear. Similar to CSA S806-12, there is a limit of 3db for the bend radius of GFRP stirrups.
ASTM D2990-09
The ASTM testing protocol for creep rupture stirrup tests calls for various figures to be plotted
in the presentation of tests for FRP materials seeking certification. The first of which deals with
plotting time to failure values for specimens at various sustained loads (Figure 2-25). From this
30
plot, one can extrapolate the design servicability stress level to be used for structures of varying
service life.
Figure 2-25 - Template for TTF graph (adapted from ASTM D2990-09)
Figure2-26, is an example of a creep modulus plot. Specimen moduli are measured and
recorded over their test life while subjected to sustained loads. As the FRP material continues
to creep until its rupture, the moduli decreases due to fibre rupturing. From the plot one can
project the design moduli that should be used in deflection analysis of structures.
Figure 2-26 - Template for modulus degradation (adapted from ASTM D2990-09)
31
Technical Information for Schöck ComBAR (Schöck GmbH 2013)
The ComBAR technical document as found on the official Schöck website provides in depth
information about their reinforcing bar products. Information from manufacturing to design
specifications to materials handling is provided. The relevant design information has been
summarized in Table 2-4. Complete material specifications of all bars from both manufacturers
are provided in Appendix B.
Table 2-4 - Material Specification (Schöck GmbH)
Specifications Straight Bar Bent Bar (12mm)
Ultimate Tensile Strength >1000 MPa 700 MPa
1000 hr 40ºC Strength* 950 MPa Not Provided
Characteristic Tensile Strength 580 MPa 250 MPa
Design Tensile Strength 445 MPa 190 MPa
Logarithmic Temporal Slope <15% Not Provided
*The sustained stress level for a 1000 hr time to failure with an ambient temperature of 40ºC
Although the actual tensile strength varies amongst the various bars, the design values provided
by Schöck are uniform across the different diameters. The ultimate tensile strength is defined
as the coupon strength of the bars. The characteristic tensile strength is defined as the strength
for a 95% confidence interval or the mean strength minus 1.64 standard deviations. The design
tensile strength is the characteristic value divided by a partial safety factor of 1.3 (German
Standard).
The technical document provides creep rupture models and values for ComBAR straight bars
and bent bars. A temporal slope value is not provided for the bent bars due to a lack of
experimental data and assumed to be equal to that of the straight bars.
From Figure 2-27 it can be seen that an increase in ambient temperature does not change the
temporal slope, ns, of the TTF curve. Schöck defines a R10 value, which is the difference in
sustained load stress over a time to failure difference of one log decade (e.g. the difference in
sustained stress to cause creep rupture failure at 1,000 hours and 10,000 hours).
32
Figure 2-27 - Time to failure example considering exposure conditions (Schöck, 2013)
2.3 Literature Review Summary The particular pieces of literature were reviewed for their relevance to the three main
themes of this thesis, creep rupture, modulus degradation, and creep strain of GFRP under
sustained loading. The experiments described dealt with straight bars, coupon tests and
beams reinforced with GFRP stirrups. However, there is a lack of understanding of the
behaviour of GFRP stirrups as shear reinforcement in concrete structures under sustained
loading. Furthermore, the design codes reviewed, CSA S806-12 and ACI 440-12R, provide
sustained load limits for straight FRP bars, and strain limits for shear reinforcement at
ultimate limit states. They do not, however, provide direct design guidance for sustained
load limits for shear reinforcement at serviceability levels or strain limits for stirrups at the
bend location. This work aims to fill these gaps by analyzing the behaviour of GFRP
stirrups under sustained loading and making evaluations of its performance as shear
reinforcement in concrete structures.
33
3 Experimental Program
In order to understand the long-term behaviour of GFRP bent bars under sustained load, 30
uni-axial tensile specimens were constructed and tested in the structures laboratory at the
University of Toronto. Individual GFRP stirrups were cast into two concrete blocks, one at
each end of the stirrup. The blocks provided a reaction surface for the hydraulic actuators,
which applied a tensile load to the GFRP stirrups. The main variables in this study were the
levels of applied tensile load, and the GFRP type. A few tests were subjected to variable loading
in order to apply the cumulative weakening model (Franke and Meyer, 1992). This chapter
describes the material properties, construction of specimens, instrumentation, equipment and
testing setup.
A list of all specimens along with variables has been provided in Table 3-1. Any properties not
shown in the table, such as concrete strength, were non-variable. The specimens have been
listed in order of descending load value. The assigned load value to specimen was selected at
random in order to avoid any unintended bias.
34
Table 3-1- List of specimens variables and testing protocol
Variable
Manufacturer
Bar
Surface
and Type
Diameter
(mm)
Bend
Radius
(mm)
Loading
Sustained
Stress
(MPa) Specimen
P3
Pultrall Sanded
(P) 12.7 53.4
Instant N/A P4
P10
P1
Sustained
383
P5 308
P7 307
P2 298
P9 288
P6 283
P8 282
Pt3
Pultrall Sanded
Twist (Pt) 12.7 83.4
Instant N/A Pt4
Pt9
Pt2
Sustained
536
Pt1 507
Pt7 503
Pt6 496
Pt8 495
Pt10 465
Pt5 NC
S3
Schöck Deformed
(S) 12 46.5
Instant N/A S4
S5
S1
Sustained
527
S2 502
S7 488
S8 455
S6 453
S9 445
S10 443
3.1 Material Properties
For this research project, the properties of the concrete and steel were not of great importance
as it has minimal influence on the behaviour of the GFRP Stirrups. The materials simply needed
to be of sufficient strength as to not allow for the specimen to undergo premature failure.
35
3.1.1 Concrete
Concrete with 28-day strength of 40 MPa, slump of 100 mm and maximum aggregate size of
14 mm was specified. The measured slump of the delivered concrete was 120 mm and there
were no workability issues or concerns. Concrete cylinders, 100 (diameter) x 200 (length) mm,
were cast in order to determine the compressive strength of the concrete (Figure 3-1). The
stress-strain characteristics of the concrete were not of much interest, only the fact that
sufficient strength had been reached prior to loading the GFRP Stirrups. Concrete strength
values vs. days after cast, has been provided in Figure 3-2. A nominal strength of 65 MPa was
reached at 28 days, the beginning of stirrup testing.
Figure 3-1 - Cylinder during testing (left), failed 100 mm diameter concrete cylinder (right)
36
Figure 3-2 - Concrete strength at 7, 28 and 62 days
3.1.2 Reinforcing Steel
Upon conducting analysis of the proposed specimen design, it was determined that the concrete
block was critical in flexure. In order to prevent premature failure by tension splitting, and
invalidating the test, it was decided to use steel reinforcement. Steel cages, pictured in Figure
3-3, not only provided the necessary strength to resist the forces of the hydraulic actuator, but
also provided mounting points for lifting hooks. Two hoops made of 12M bars were welded
together at various points, with straight bars (Figure 3-3 right). All steel used was specified as
400W.
0
10
20
30
40
50
60
70
0 10 20 30 40 50 60 70
Stre
ngt
h (
MPa
)
Time (Day)
37
Figure 3-3 - Steel reinforcement of specimens
3.1.3 Glass Fibre Reinforcing Polymers Stirrups
Pultrall Inc. (Company A) provided twenty stirrups of two types, and Schöck (Company B)
provided ten stirrups of a single type. All stirrup bars had a nominal diameter of 12 mm,
however, the measured diameter differed. Nominal core diameter and outside diameter and
their respective areas are provided in Table 3-2. For the purposes of analysis, the inside
diameter was used. The discrepancy between the inside and outside diameter is due to the
coating applied to the bars, the Pultrall bars had a sand coating of 0.25 mm thickness, Figure
3-4, while the Schöck bars had a plastic sleeve of 0.5 mm thickness.
Table 3-2 - GFRP Stirrup dimensions
Bar Type Nominal
Diameter
(mm)
Nominal
Area
(mm2)
Outside
Diameter
(mm)
Outside
Area
(mm2)
Inside
Diameter
(mm)
Inside
Area
(mm2)
P 12.7 127 14 154 13.5 143
Pt 12.7 127 14 154 13.5 143
S 12 113 13 133 12.0 113
38
Figure 3-4 - Inside and outside diameters for sand coated and plastic coated bars
(Johnson, 2014)
In order to establish the straight bar strength of the GFRP, coupons were especially ordered
from the manufacturers and tested using a 1000 kN capacity testing machine (MTS 311.31).
Tests were conducted in accordance with Annex C of CSA S806-12.
Clamping directly onto the GFRP bar would crush the fibres and render the tests invalid.
Therefore, couplers were used in order to test the GFRP bars. HSS sections with a diameter of
approximately 50 mm (2 “) were used for couplers. Rock Frac, an expansive mortar, in the
space between the bar and inside of the coupler provided the clamping force to hold the GFRP
bars. The coupler system with the Rock Frac and polymer o-rings for sealing are displayed in
Figure 3-5. Three tests for each bar type were carried out.
Figure 3-5 - Tensile coupon tests with steel couplers
39
Figure 3-6 - External MTS gauge
Figure 3-7 - Coupon after testing, failure in-between couplers
An MTS displacement gauge was used to determine the stress-strain response and modulus of
elasticity of the bar being tested (Figure 3-6). Failures in all bars occurred in-between couplers
(Figure 3-7). Stress strain profiles have been provided in Figures 3-8 through 3-10 for the three
types of bars. A summary of the test results are provided in Table 3-3 including the calculated
moduli. The stress presented is based on the effective/inside measured diameter. The MTS
gauge was removed at about 60% of the ultimate stress to avoid damage and the response
beyond that point is presented by assuming a linear behaviour.
40
Figure 3-8 - Deformed 12 mm coupon (S) test results
Figure 3-9 - Twisted 12 mm coupon (Pt) test results
0
200
400
600
800
1000
1200
0 0.005 0.01 0.015 0.02
Ten
sile
Str
ess
(MPa
)
Strain (mm/mm)
0
200
400
600
800
1000
1200
1400
0 0.005 0.01 0.015 0.02 0.025
Ten
sile
Str
ess
(MPa
)
Strain (mm/mm)
41
Figure 3-10 - Sanded 12 mm coupon (P) test results
It can be seen that there exists some significant variability in the ultimate strength for the Pt
series and to a lesser degree for the S Series. This is important due to the fact that variability
of the strength would further increase when the bar is bent, as the specimen would undergo
significant non repeatable damage. Voids and buckling of fibres would be created in a non-
uniform nor predictable fashion.
Table 3-3 - Coupon test summary and material strength
Specimen
Ultimate
Load
(kN)
Avg. Ult.
Load
(kN)
Avg. Ult.
Stress
(MPa)
Avg. Modulus
(MPa)
Pt-1s 162.22
150.9 1054 58800 Pt-2s 137.97
Pt-3s 152.51
P-1s 155.8
157.4 1100 52490 P-2s 158.96
P-3s 157.52
S-1s 116.9
112.1 992 59800 S-2s 112.7
S-3s 106.7
The ultimate tensile load of the Pt and P series is higher, as it has a larger actual load bearing
area (143 mm2) compared with that of the S series (113 mm2). The nominal areas for bars P,
Pt and S are given to be 127, 127 and 113 mm2, respectively. With these differences in area
0
200
400
600
800
1000
1200
1400
0.000 0.005 0.010 0.015 0.020 0.025
Stre
ss (
MPa
)
Strain (mm/mm)
42
taken into consideration, the strengths of the bars are nearly equal. Although the S Series bars
have a larger modulus, an important design factor, the P Series bars were able to undergo a
larger strain.
The stirrups used for this project were all specified to have the minimum allowable inside bend
radius of 3db as per CSA-S806 with a bend radius no larger than 4db. The dimensions of the
stirrups which were submitted for fabrication are provided in Figure 3-11 and pictures of the
stirrups are shown in Figure 3-12. The reason for the lower limit was to test specimens that
were allowable under current codes therefore making the test results more relevant and
reasonably conservative. The upper limit was put in place to safeguard against specimens that
would be too strong and exceed the capacity of our load cells and hydraulic equipment. Figure
3-13 shows the bend radius calculation, the measurement was taken to the inside of the bar. A
summary of the bend radii has been provided in Table 3-4. It is evident from Figure 3-12 that
the Pt Stirrup does not have a clear start of the bend section. In this case, the radius was
calculated using software (BlueBeam) by selecting three points along on the bend (Figure 3-
14). The bend radius to diameter ratio is calculated using the effective diameter in order to
make use of the JSCE equation (equation 2-5).
Figure 3-11 - GFRP Specimen dimensions as ordered from suppliers
43
Figure 3-12 - From left to right – P (left), Pt (centre), and S (right) Stirrups
Table 3-4 - Bend radius properties of GFRP stirrups
Stirrup type Bend
Radius
Bend Radius/
Diameter Ratio
P 47 3.96
Pt 66 4.88
S 40 3.33
Figure 3-13 - Bend diameter for P (Left) and S (Right) stirrups
44
Figure 3-14 - 3 Point bend radius of Pt stirrup
The bend region for all three stirrups was milled out in order to see the fibre distribution. The
P Series had moderately uniform fibre distribution but had damage and higher amounts of resin
in the inside of the bend (Figure 3-15). The Pt Series had the highest degree of fibre non-
collinearity but uniform fibre distribution (Figure 3-16). The S Series had high level of
uniformity of fibres, and low degree of kinking (Figure 3-17).
45
Figure 3-15 - Internal fibres of P stirrup bend region
Figure 3-16 - Internal fibres of Pt stirrup bend region
46
Figure 3-17 - Internal fibres of S stirrup bend region
3.2 Construction of Specimens
The formwork consisted simply of two boxes, with inside dimensions of 500 x 200 x 200 mm,
constructed of 19 mm thick Plywood. The two boxes were joined together with “2x4’s” which
maintained the parallelism of the two reaction surfaces, and provided carrying points, pictured
in Figure 3-18.
47
Figure 3-18 - 3D View of formwork and specimen
3.3 Instrumentation and Testing Equipment
Due to the nature of the long term tests, it was paramount that this project should not tie up
important lab resources which were required by other projects. Therefore, all equipment
acquired was specifically for this project in order to remain independent. New load cells,
pumps, hydraulic jacks, hoses, pressure switches, data acquisition systems and computers were
required. The testing equipment and instrumentation directly relevant to the research project is
briefly explained in the following sections.
3.3.1 Strain Gauges
In order to capture the strain behaviour and make it possible to predict rupture, strain gauging
of the specimen was required. Strain gauges along the straight portion were required to confirm
concentric loading and determine the straight portion moduli versus bend region moduli. Five
strain gauges were placed along the bend section to capture where a potential creep rupture
48
failure occurred, along with the stress distributions. A schematic showing the location of all
eight strain gauges is provided in Figure 3-19.
Figure 3-19 - Strain gauge locations (typical)
Due to the long duration of the testing, special strain gauges needed to be used. Typically,
gauges with a resistance of 120 ohm’s is preferred, however, the voltage excitation required
could have caused voltage saturation for tests lasting several days let alone months. Therefore,
350 ohm gauges were used in order to reduce the voltage input and still achieve the required
resolution for strain gauge measurements. For additional information about the strain gauges
used, refer to the product summary sheet provided in the Appendix B.
49
Figure 3-20 - Removal of sand coating and placement of 2mm gauge
Figure 3-21 - Gauge placement on bend, 4 and 5, and wrapping of foil tape
The 350 Ohm, 2 mm, gauges were placed onto the specimen using the procedures outlined in
the product package and best practice steps established from experience. Firstly, the exterior
surface, sand or plastic sleeve, was removed in order to allow placement of the gauge on the
load bearing surface (Figure 3-20). The bar was then cleaned with acetone, and a strain gauge
affixed using the CN Glue. This was followed by a layer of M-Coat. Afterwards the strain
gauge was covered in protective SB tape and water proofing foil tape (Figure 3-21). Finally
50
the region which required to be de-bonded from the concrete was wrapped in a layer of tape
as to prevent reinforcement development.
3.3.2 Load Cells
The maximum expected tensile capacity of the GFRP stirrups was expected to be about 150
kN with the sustained load levels being lower. Honeywell Model 53 Load Cells with a 133 kN
capacity (67 kN additional reserve) were used for testing, pictured in Figure 3-22. A product
sheet has been attached in the Appendix B, which provides additional details such as
dimensions and circuitry. Figures 3-23 and 3-24 are of a specimen prior to being cast and
casting taking place for all 30 specimens, respectively.
Figure 3-22 - Honeywell Model 53 Load Cell
51
Figure 3-23 - Formwork with specimen prior to cast
Figure 3-24 - Casting of all 30 specimens
3.3.3 Hydraulic Jacks
In order to apply the required tensile load of 150 kN +, hydraulic actuators with a capacity of
200 kN (20 tonnes) were acquired (Figure 3-25). The hydraulic cylinders were placed in
between the two concrete blocks and powered using a pneumatic hydraulic pump. The lab air
supply of 1.4 MPa (200 psi), was converted to a hydraulic pressure of 67 MPa (10,000 psi) for
use in the jacks. Four hydraulic jacks were connected to one pump through a system of
hydraulic hoses running in parallel.
52
Figure 3-25 - 200 kN hydraulic cylinders
A maximum of eight tests ran simultaneously in order to complete the testing program in a
reasonable time. Parallel testing allowed for the staggering of tests, 100 day tests were started
first using half of the eight test rigs, while the shorter term tests were cycled through the
remaining four test rigs. The test setup, with all instrumentation can be seen in Figure 3-26. A
pressure switch with a solenoid was used, pictured in Figure 3-27, to control the pressure.
Figure 3-26 - Test specimen with hydraulic cylinder, and potentiometers
53
Figure 3-27 - Pneumatic to hydraulic pump with pressure switch
3.4 Test Observations
This section discusses some of the observations made during the testing phase of the project in
order to provide qualitative insight to complement the quantitative analysis which follows in
Chapter 4.
The testing procedure called for eight tests to be run in parallel, powered by two hydraulic
pumps. It was intended that the four tests all connected to the same hydraulic line would be
subject to the same loading. However, this was not the case. Although the same pressure was
present in the lines for all four tests, the load cells recorded varying loads. Initially it was
thought the issue was caused by the difference in the manufacturing of the individual hydraulic
jacks or load cells. In an attempt to remedy this, load cells and hydraulic jacks were labelled to
keep track of their biases, and tested independently. No pattern or bias was found for any of
the equipment. The reason for the discrepancy, albeit small (1-3%), remains unsolved.
To ensure loading be applied equally and uni-axially, the hydraulic actuators and load cells
were placed at the midpoint between the left and right straight portions of the GFRP. Upon
loading, strain gauge data between 1, 2, and 8, which should be nearly identical, varied.
54
Adjustments to the position of the hydraulic actuator were made to counter act the strain values
(i.e. if gauge 2 read a higher value than gauge 8, the actuator was moved closer to gauge 8).
Adjustment of the actuator yielded some improvement but complete uniformity could not been
reached. Upon closer inspection, it was discovered that the GFRP bars were not completely
straight. Nearly all bars, some more than others, possessed bowing. Strain gauges placed on
the convex side showed a smaller strain than those on the concave side. As loading was
increased and the GFRP bar straightened, the concave side witnessed higher elongation than
its opposite.
During the first test it was discovered that the specimens were exhibiting unwanted out of plane
bending. A steel frame was constructed in order to prevent the rotation of the specimens, and
force the bars to undergo only uniaxial tensile loading. A specimen in its steel frame is pictured
in Figure 3-28. This frame solved the issue of the rotation and all the three strain gauges
mounted on the straight portions and potentiometers measured equal strain values.
Figure 3-28 - Specimens under sustained load inside steel frame
55
4 Analysis of Data
Thirty tests were conducted on GFRP stirrups, ten of each type (S, P, and Pt). Of the ten tests
conducted for each type, three were under instant loading to failure, 100% tests, used to
establish the ultimate strength of the stirrup. Additionally, nine straight bar coupon tests were
conducted, three of each type, used to determine the bend strength to straight bar strength ratio.
4.1 Stirrup 100% Tests
The ultimate load tests were conducted by monotonically increasing the load approximately
0.2 kN/sec, with the average test spanning about 10 minutes. However, as our loading
mechanism was pneumatically powered and pressure controlled, there was some variability in
the loading rate. Figures 4-1 to 4-3 show applied load vs. strain behaviour for specimens P3,
Pt9 and S4 respectively. These specimens represent one from each series. Figures 4-4 to 4-7
show typical failures where the final stages of specimens P3, Pt9, Pt2 and S4 are presented,
respectively.
Figure 4-1 - Specimen P3 - Strain and Applied Load vs. Time
0
20
40
60
80
100
120
0
2000
4000
6000
8000
10000
12000
14000
16000
150 200 250 300 350 400 450 500 550 600 650
Ap
plie
d L
oad
(kN
)
Stra
in (
um
/m)
Time (seconds)
Gauge 1
Gauge 2
Gauge 3
Gauge 4
Gauge 6
Gauge 8
Load
56
Figure 4-2 - Specimen Pt9 - Strain and Applied Load vs. Time
Figure 4-3 - Specimen S4 - Strain and Applied Load vs. Time
0
20
40
60
80
100
120
140
160
180
0
2000
4000
6000
8000
10000
12000
14000
16000
18000
0 100 200 300 400 500 600 700
Ap
plie
d L
oad
(kN
)
Stra
in (
um
/m)
Time (seconds)
Gauge 1
Gauge 2
Gauge 3
Gauge 4
Gauge 5
Gauge 6
Gauge 7
Gauge 8
Load
0
20
40
60
80
100
120
140
160
180
0
5000
10000
15000
20000
25000
0 50 100 150 200 250 300 350 400
Ap
plie
d L
oad
(kN
)
Stra
in (
um
/m)
Time (seconds)
Gauge 1
Gauge 2
Gauge 3
Gauge 4
Gauge 5
Gauge 6
Gauge 7
Gauge 8
Load
57
Figure 4-4 - Typical failure of P Series specimens at bend location (P3 pictured)
Figure 4-5 - Pt 9 specimen failure at bend location
58
Figure 4-6 - Pt 2 after failure with bend pull out
Figure 4-7 - Typical failure of S Series specimens at bend location (S4 pictured)
The nine stirrup tests under instant loading have been summarized in Table 4-1, along with
relevant information such as bend strength retention and predicted values according to the
JSCE equation. The S Series was able to maintain the highest level of strength (fbend/fstr = 0.67)
while the P series with the lowest (fbend/fstr = 0.36). This may be caused by the fact the P Series
suffered from an uneven distribution of resin and fibres (Figure 3-15). Although the nominal
bend radius to diameter ratio of the Pt series provided by the manufacturer is 4, the effective
59
radius was 4.88 times the diameter of the bar. The bend strength retention values are compared
to those obtained from JSCE equation 2-5 using the nominal bend radius.
Note the variability of the ultimate strength of bends is much greater than that of the straight
bars. Again, this is due to the randomness of the kinking of fibers and voids formed during the
bending of the bar. This raises concerns for the consistency and reliability of the FRP bent bars
and their use in structures. Premature failure and variability is discussed in later sections.
From Table 4-1, it can be seen that the JSCE equation accurately predicted the bend strength
to straight bar value for Pt and S Stirrups. The fbend/fstraight values coincide with the test results
obtained by Ahmed, El-Sayad, El-Salakawy, and Benmokrane (2010).
Table 4-1 - Summary of bend strength properties
Specimen
Ultimate
Load
(kN)
Avg.
Ult.
Load
(kN)
Avg. Ult.
Stress
(MPa)
fbend/fstraight
JSCE
Prediction of
fbend/fstraight
Actual/
JSCE
Prediction
Pt-3 175.3
161.5 564 0.54 0.54 0.99 Pt-4 153.0
Pt-9 156.3
P-3 100.0
113.5 397 0.36 0.48 0.75 P-4 118.0
P-10 122.6
S-3 161.0
151.1 668 0.67 0.47 1.42 S-4 163.3
S-5 132.2
All specimens failed at the instrumented bends, and no failures occurred in the fully bonded
side, see Figure 4-8. Failures typically occurred, at what appears to be the location of gauge 4
and 6. This corresponds to the higher strain values generally observed at those locations.
60
Figure 4-8 - No damage or failures at other fully bonded end (typical)
.
4.1.1 Time to failure (TTF) for monotonic loading
In order to use the 100% load tests as data points in the calculation of TTF charts, a sustained
load time is needed to be assigned. Equation 2-4 was used to determine the equivalent time
(t1). Equation 2-4, is provided below along with the relevant values of test Pt3 inserted. The
value of tb was taken as the test duration and value of σ1 was taken as the failure stress (615
MPa). The value of ns was determined through iteration; initially, a value of ns = 1 was used
for all three Series, and after two iterations convergence was reached. The overall TTF
regression was not very sensitive to the value of ns for the 100% tests. A summary of the TTF
values has been provided along with the test duration in Table 4-2. The process used to
calculate the equivalent time and the cumulative weakening model is explained in detail in
Chapter 4.5.
𝑡1 =1
𝜎1𝑛𝑠 ∫ 𝜎(𝑡)𝑛𝑠𝑑𝑡
𝑡𝑏0
[2-4]
𝑡1 =1
61510.98∫ 𝜎(𝑡)10.98𝑑𝑡1258
0
61
Table 4-2 - Monotonic loading test duration and equivalent time
Specimen Ultimate
Load
(kN)
Test
Duration
(sec)
ns Equivalent
Sustained Load
Time (sec)
Pt3 175.3 1258 10.98
145.9
Pt4 153.0 571 40.3
Pt9 156.3 679 66.0
P3 100.0 599 10.64
37.3
P4 118.0 320 39.4
P10 122.6 350 34.7
S3 161.0 325 20.50
11
S4 163.3 376 36.2
S5 129.0 437 35.2
62
4.2 Creep Rupture Results
Table 4-3 - Time to failure summary for sustained load tests
Specimen Load (kN) %Ult TtF (s) TtF (hr)
Pt2 142.2 88 1160 0.3
Pt6 131.6 81 157000 43.6
Pt7 133.6 83 159000 44.2
Pt10 123.4 76 186500 51.8
Pt1 134.5 83 1146700 319
Pt8 131.5 81 1275000 354
P1 101.8 93 8300 2.3
P7 81.5 75 10350 2.9
P5 81.8 75 12500 3.5
P6 75.2 69 2218800 616
P8 74.9 69 2849400 792
P9 76.3 70 4700000 1306
P2 79.0 72 9100000 2530
S7* 112.1 74 307 0.1
S2 113.6 75 830 0.2
S10 100.3 66 30000 8.3
S1 119.3 79 41600 11.6
S9 100.6 67 963600 263
S6 103.0 68 1630800 453
S8 103.0 68 3713700 1032
*Specimen S7 was found to be defective and not considered in the analysis
Duration of the twenty-one GFRP stirrups tests varied from a few minutes to over 2500 hours
(105 days) with stress levels ranging from 93% to 66% of the average ultimate instant strength.
Furthermore, a failure occurred in every time epoch (1 hr, 10 hr, 100 hr, and 1000 hr). All tests,
ultimate and sustained, of all three types have been plotted in Figure 4-9. The time to failure
63
is provided in seconds, on a logarithmic scale (X-axis). The load is provided as a percentage,
all loads were divided by the average value of ultimate instant load capacity for their respective
type (e.g. S1’s sustained load of 119.3 kN is divided by 151.1 kN giving a value of 79% of
ultimate). Overall, the coefficient of determination, R2, is moderately strong for all three types,
with the P series being of the highest strength of 0.82 and the Pt and S series having regression
strength of 0.69 and 0.73, respectively.
Figure 4-9 - Applied load vs. Time to Failure for S, P and Pt Series
0%
20%
40%
60%
80%
100%
120%
1.E-04 1.E-03 1.E-02 1.E-01 1.E+00 1.E+01 1.E+02 1.E+03 1.E+04 1.E+05
Perc
enta
ge o
f U
ltim
ate
Load
Time to Failure (hr)
S Series
P Series
Pt Series
64
4.2.1 S Series
Figure 4-10 - S-Series - Applied Stress vs. Time to Failure
The S Series’ sustained load tests ranged from 15 minutes to 1031 hours, covering four time
epochs, see Figure 4-10 for TTF chart. Typically, failures occurred at the end of the bend
region, at strain gauge location 6. The average strain upon failure was 1.57 % while the average
corresponding straight portion strain was 1.1% (see Table 4-4). Furthermore, the strains at other
parts of the bend, gauges 3, 4, 5 and 7, ranged from 0.8% to 1.2%.The failure at location 6 with
the higher strain value, up to 60% greater, may be explained by Figure 2-16. As this region is
not only subject to a tensile load but a significant shear load.
Table 4-4 - S- Series strain values at failure locations
Strain (um/m)
Specimen S1 S2 S6 S7 S8 S9 S10 Average
Gauge 1 14100 9600 9400 8000 9800 9500 8800 9900
Gauge 2 9400 10500 8600 10900 7400 8800 11700 9600
Gauge 3 N/A 12200 6300 10200 3600 N/A 6200 7700
Gauge 4 N/A 14500 8300 N/A 580 5300 N/A 8500
Gauge 5 N/A 14300 N/A N/A 14900 7900 N/A 12300
Gauge 6 17700 18100 N/A 14100 14400 12600 17500 15700
Gauge 7 N/A N/A 3500 6700 10000 8600 N/A 7200
Gauge 8 14900 12700 N/A 13100 11700 9800 7500 11600
y = -19.25ln(x) + 553.73R² = 0.73
0
100
200
300
400
500
600
700
800
0.00 0.01 0.10 1.00 10.00 100.00 1000.00 10000.00
Ap
plie
d S
tres
s (M
Pa)
Time to Failure (hr)
65
4.2.2 Pt Series
Figure 4-11 - Pt Series, Applied Stress vs. Time to Failure
The Pt Series’ sustained load tests ranged from 20 minutes to 354 hours, covering three time
epochs (Figure 4-11). Failures typically occurred in the end of the bend (Gauge 6) and some at
the location of splicing between the straight and bend portions (Gauge 3). The exact location
of the failure is not always known, but it was deduced from the strain data (Table 4-5) and
visual evidence from the bend pull out. The stirrup elongation at failure at location 1 through
8, excluding 5, 7, were about 1.1%. The general uniformity of the strains may be attributed to
the fact the Pt Series stirrup has a larger bend radius than the P and S Series.
Table 4-5 - Pt Series strain values at failure locations
Strain (um/m)
Specimen Pt1 Pt2 Pt6 Pt7 Pt8 Pt10 Average
Gauge 1 9300 11000 10000 N/A 8700 10300 9900
Gauge 2 9500 10600 10700 8100 8300 9700 9500
Gauge 3 13600 10500 9200 8700 10500 12100 10800
Gauge 4 13000 N/A 12900 12200 11000 11400 12100
Gauge 5 N/A N/A 7700 6600 N/A N/A 7200
Gauge 6 9900 N/A 8100 12700 10900 11200 10600
Gauge 7 N/A 5800 N/A N/A N/A N/A 5800
Gauge 8 9200 10700 12500 9800 9200 7600 9900
y = -11.90ln(x) + 549.98R² = 0.69
0
100
200
300
400
500
600
700
0.01 0.10 1.00 10.00 100.00 1000.00
Ap
plie
d S
tres
s (M
Pa)
Time to Failure (hr)
66
4.2.3 P Series
Figure 4-12 - P Series, Applied Load vs. Time to Failure
The P Series’ sustained load tests ranged from 2.3 hours to 2528 hours, covering four time
epochs (see Figure 4-12). The average elongation at locations of failure was 1% while the
corresponding straight portion stress was 0.75% (see Table 4-6). Furthermore, the strains at
other parts of the bend, gauges 3, 4, 5 and 7, were also 0.75%. The failure at location 6 and
coinciding with high strain value, 33% greater, may again be caused by the shear stress present
at that location.
Table 4-6 - P Series Strain values at failure locations
Strain (um/m)
Specimen P1 P2 P5 P6 P7 P8 P9 Average
Gauge 1 8600 6500 5800 5400 5900 7700 5300 6500
Gauge 2 8700 7000 7900 7300 5100 11200 6200 7600
Gauge 3 6900 4700 7400 13500 11600 9500 6800 8600
Gauge 4 8800 5600 6200 N/A 3700 N/A 5700 6000
Gauge 5 N/A 3600 N/A 10800 N/A N/A N/A 7200
Gauge 6 5900 16600 8600 7000 10600 11600 10200 10100
Gauge 7 4300 3000 4500 9300 N/A 11600 1510 5700
Gauge 8 6600 6500 8800 5800 6900 10700 4800 7200
y = -11.01ln(x) + 358.91R² = 0.82
0
50
100
150
200
250
300
350
400
450
500
0.00 0.00 0.01 0.10 1.00 10.00 100.00 1000.00 10000.00
Ap
plie
d S
tres
s (M
Pa)
Time to Failure (hr)
67
4.3 Creep Rupture and Sustained Loads
One of the key objectives of this research project is to find the “safe” sustained stress level for
GFRP stirrups with a lifespan of 50 to 100 years. Using a simple logarithmic extrapolation,
one is able to calculate the applied stress to cause failure at 50 years and 100 years. The fitted
equations, with projected strengths have been provided in Table 4-7, where x is in hours, and
y is equal to the sustained stress in MPa. A time value greater than 1 hour should be used for
all three equations as test data shorter than this duration is limited. The 50 year and 100 year
sustained load capacities have been calculated, along with the percent of the average ultimate
stress the specimen would be able to sustain over that time. Note the insignificant change in
load level, from 50 to 100 years. In this time domain, the loads are simply too low to cause any
direct structural damage of GFRP. Failure would likely only occur by the slow diffusion of
alkali or water into the bar, causing delamination or degradation of fibres.
The slope of the TTF curve allows one to infer the materials resilience to creep rupture. A slope
of near zero would indicate the specimen is not significantly affected by sustained loading (e.g.
steel), a lower number would signify a higher creep stress compliance. Although the S Series
maintains significant bend strength, it is the more susceptible to creep stress than the other two
bars. This may be due to the fact it has a different manufacturing process and material
properties. The P and Pt series stirrups are not formed continuously, but a straight bar is joined
together with a bend bar using epoxies. While the S Series is one bar that is bent and heat cured.
It should be noted that X=1 in these equations will provide a one-hour sustained stress capacity.
Table 4-7 - Time to failure regression equations
Series Sustained stress as a function
of time to failure
50 Year Capacity 100 Year Capacity
Pt y = -10.98ln(x) + 510.88 368 MPa (65%) 359 MPa (63%)
P y = -10.64ln(x) + 338.37 200 MPa (50%) 193 MPa (49%)
S y = -20.50ln(x) + 550.48 284 MPa (42%) 270 MPa (40%)
Recall Weber (2005) conducted similar creep rupture tests on Straight ComBARs to produce
similar Time to Failure charts. Schöck’s tests were conducted at various ambient temperatures,
30ºC, 40ºC and 60ºC. The ambient temperature for the lab in which our tests were conducted
was roughly 22ºC. The larger creep stress compliance coefficient, -64.9 of the straight bar, is
much larger in magnitude than the creep stress compliance coefficient of -20.59 for the S
68
Series. The reason for this is not clear. A variance in temperature, bar size, or environment
should not change the logarithmic temporal slope, only the y intercept, according to research
by Schöck (2013). The R10 value provided by Schöck (2013) for GFRP is 15% and for CFRP
5%, indicating the material composition to be the main factor in the magnitude of the slope. It
is therefore possible that S Series R10 value of 15% is due to a difference in the material
composition between the straight and bent bars. Recall, a R10 value is the difference in
sustained load to cause failure over a log decade (a multiple of 10). Furthermore, the sustained
load of 40% of URS to cause failure at 100 years is similar to the findings by Weber (2005)
who had a value of 46% of URS.
Chapter 2.0 summarizes the testing conducted by Johnson and Sheikh (2014) regarding the
analysis of beams reinforced with GFRP shear reinforcement. Johnson’s thesis (2014) provides
shear reinforcement strain data at 450 kN which approximately corresponds to SLS conditions.
Clause 8.4.5.2 from S-806-12 stipulates a minimum shear reinforcement ratio of 0.35% (f’c =
35 MPa), approximately bisecting 0.22% and 0.40% reinforcement ratio provided. Equation 9-
7 from ACI 440.1R-12 stipulates a minimum shear reinforcement ratio of 0.22% (for GFRP
shear reinforcement strength ffv = 160 MPa). By using the bend location strains from Figure 4-
13 we can calculate the expected TTF values by using the respective equations found in Table
4-7 (see Table 4-8 for TTF values). Figure 4-14 pictures the typical bend strain gauge location
on GFRP stirrups used.
Figure 4-13 - Stirrup strain values for beams tested by Johnson and Sheikh (2014)
69
Figure 4-14 - Location of strain gauge for tests conducted by Johnson and Sheikh (2014)
Table 4-8 - Time to failure projections for beams (Johnson and Sheikh, 2014) under SLS (450
kN) loading
Specimen
Peak Bend
Location Strain
(mm/m) Stress (MPa)
Time to
Failure
(Yrs.)
JSC32-22B 0.0025 157.5 24100
JSC32-40B 0.0032 201.6 2810
JSC32-50B 0.0025 157.5 24100
JSV40-22B 0.0048 254.4 0.30
JSV40-40B 0.0038 201.4 44.5
JSV40-50B 0.0011 58.3 30800000
JSH40-50B 0.001 63 2400000
JSC100-50B 0.0048 302.4 20.6
The expected TTF values, from the shear critical beams tested, only three have values which
are of concern, JSV40-22B, JSV40-40B, and JSC100-50B. The first has a shear reinforcement
which is unacceptable under current CSA standards. The second has a fairly high TTF value
of 44.5 years. The third beam is of great concern, not only is the TTF value below the typical
service life but also, the shear reinforcement of 0.5% is significantly above the 0.33%
minimum.
One issue with the projected time to failures shown in Table 4-8 is that the failure mechanism
is not deterministic. There is an inherent variance in the material as previously discussed. This
70
variance is accounted for by Schöck (2013), in which the design values for GFRP have been
reduced by one time epoch and multiplied by environmental factors. This is similar to that of
the ACI 440 method of applying a CE factor of 0.8 for normal exposure or 0.7 for weather and
earth exposure. Franke and Meyer (1992), recommended the use of a safety factor based on the
average strength minus three standard deviations.
If we look at the difference between time to failure and projected values for the tests conducted,
we can get a sense of variability. Consider the ratio of Actual TTF (TA) divided by the Predicted
TTF (TP), see Figure 4-13 for clarification. The TA/TP ratio has been plotted for all 18
specimens in Figure 4-16.
Figure 4-15 - TA/TP illustration
71
Figure 4-16 - TA/TP values for all specimens
A TA/TP value of one means the specimen failed when predicted, while a value of 10 means
the specimen failed one log decade later than expected. If we find the standard deviation of all
TA/TP values for each series and subtract them from the expected TTF values from Table 4-8
we obtain a quantitative idea of a premature TTF values (Table 4-9).
Table 4-9 - Premature time to failure
Series Bend Stress
(MPa)
Nominal TTF
(hours)
TTF - 1 STD
(hours)
TTF - 2 STD
(hours)
TTF - 3 STD
(hours)
P 254.4 2680 136.0 6.90 0.35
S 302.4 180100 15040 1256 104.9
Once the variability of the material is taken into account, the TTF values fall to dangerously
low levels. The standard deviation “levels” have been added to the original creep rupture curves
to show the possible failure range based on our test results (Figures 4-17, 4-18 and 4-19).
Standard deviation levels of minus one, two and three have been plotted along with
corresponding bend stress levels for the P and S series (the Pt series was not tested).
0.01
0.1
1
10
100
T A/T
P
72
Figure 4-17 - P Series Applied Load vs. TTF with 1, 2 and 3 standard deviation levels
Figure 4-18- Pt Series Applied Load vs. TTF with 1, 2 and 3 standard deviation levels
1 STD
2 STD
3 STD
JSV40-22B Bend Stress Level
0
100
200
300
400
500
1.E-03 1.E-02 1.E-01 1.E+00 1.E+01 1.E+02 1.E+03 1.E+04 1.E+05 1.E+06
Ap
plie
d S
tres
s (M
Pa)
Time (hours)
2 STD
1 STD
3 STD
0.0
100.0
200.0
300.0
400.0
500.0
600.0
700.0
1.E-03 1.E-02 1.E-01 1.E+00 1.E+01 1.E+02 1.E+03 1.E+04 1.E+05 1.E+06
Ap
plie
d S
tres
s (M
Pa)
Time (hours)
73
Figure 4-19 - S Series Applied Load vs. TTF with 1, 2 and 3 standard deviation levels
The shear stress for a rectangular section varies parabolically with the greatest value at the
neutral axis (Figure 4-20). The bend in a typical stirrup ended at approximately 100 mm from
the bottom of the beam. The shear stress at that location is about 55% of the shear at the mid-
section (τmax). This may lead to the conclusion the weaker bend section is not of concern as the
stress is only 55% of the mid-depth stress. However, as Table 4-1 summarizes, the bend
strength may be nearly half that of the straight portion and therefore under the same tensile
demand as the straight portion. Furthermore, as the composite section cracks, the concrete
contribution would drop and the shear reinforcement contribution would increase. This is
evidenced by the fact that seven out of eight beams presented in Figure 4-13 have almost the
same strain values for the bend and straight locations.
2 STD
1 STD
3 STD
JSC100-50B Bend Stress Level
0
100
200
300
400
500
600
700
800
1.E-03 1.E-02 1.E-01 1.E+00 1.E+01 1.E+02 1.E+03 1.E+04 1.E+05 1.E+06
Ap
plie
d S
tres
s (M
Pa)
Time (hours)
74
Figure 4-20 - Shear stress distribution along rectangular cross section
4.3.1 Finite Element Modelling - VecTor 2
Beam JSV40-40B and JSC100-50B were analysed using VecTor 2 to obtain the transverse
reinforcement stresses to corroborate the data provided by Johnson and Sheikh (2014). The
model created was similar to that of Johnson and Sheikh (2014) as to keep analysis consistent.
A visual of the FEM model used has been provided in Figure 4-21. The material properties that
were used for the Finite Element Modelling (FEM) program have been summarized in Table
4-10.
Figure 4-21 - Formworks beam model
For analysis, all default models were used, except for the high strength concrete beams, where
Popovics pre and post peak behaviour model was used. The mesh size was reduced until
75
convergence was reached at 25 mm x 25 mm. Only half the beam was modelled in order to
reduce run time. Furthermore, a bearing plate was modelled using properties of steel in order
to distribute the load and prevent local failures. The model was load controlled, with 2 kN
increments, until failure was reached.
Table 4-10 - Material properties input for Formworks
Bar Cross
Sectional
Area
(mm2)
Bar
Diameter
(mm)
Yield
Strength
(MPa)
Ultimate
Strength
(MPa)
Elastic
Modulus
(MPa)
Strain
Hardening
Value (mε)
Ultimate
Strain
(mε)
Bar B1 198 16 1,264 1,264 71,000 17.8 17.8
Bar C 200 16 1,234 1,264 61,000 20.3 20.3
B1 Stirrups
(S)
147 13 655 655 41,500 15.8 15.8
B1 Stirrups
(B)
147 13 507 507 41,500 12.2 12.2
C Stirrups
(S)
95 11 912 912 57500 15.9 15.9
C Stirrups
(B)
95 11 531 531 57500 9.24 9.24
(S) Straight Portion, (B) Bend Portion
Two versions were modelled for each beam, one taking into account the lower strength of the
bend region (VecTor 2-Bend), and the second which used the strength of the straight portion
in the bend region (VecTor 2). This was done to see the effect the weaker bend would have on
the capacity of the beam at failure but also stress levels and deformation at serviceability loads.
JSV-40-40B
The analytical load deformation plots and the experimental behaviour for beam JSV-40-40B,
40 MPa concrete with 0.4% shear reinforcement using V-Rod stirrups, have been provided in
Figure 4-22. Overall the pre peak behaviour was captured adequately. The post peak was not
captured by VecTor 2. Comparing the two FE models, the response was similar at 450 kN
beyond which the two curves diverge. At failure the beam capacity from the VecTor 2 (Bend)
analysis is about 100 kN lower than that from VecTor 2. However, the actual beam had a shear
failure at mid-depth, see Figure 4-23, implying that the bend region was not as critical in this
particular beam. The failure load from the model without bend consideration was found to be
similar to the experimental capacity.
76
Figure 4-22 - Load vs. Midspan Deflection responses from VecTor 2 models and experiment
Figure 4-23 - Failure region of JSV-40-40B (Johnson and Sheikh, 2014)
The reinforcement tensile stress for approximately service load conditions output from
Augustus has been provided in Figure 4-24. The maximum shear reinforcement stress at mid-
depth is equal to 235 MPa, with the bend region stress 161 MPa. A stress level of 235 MPa
corresponds to a strain of 0.0057 which is about 40% greater than the ACI strain limit of 0.004.
The peak bend stress reached for JSV-40-40B was 201 MPa (0.0048 mm/mm). This is of
concern, because the mid span strain is higher and above the ACI limit, causing the section to
0
100
200
300
400
500
600
700
800
900
1000
0 10 20 30 40 50
Load
(kN
)
Midpan Deflection (mm)
VecTor 2
Experiment
VecTor 2 (Bend)
77
be deficient in two respects. Deficient aggregate interlock at mid-depth and an over stressed
bend region.
Figure 4-24 - Augustus shear reinforcement stress at SLS (450 kN)
A vertical slice of the beam was taken at quarter span to determine the shear stresses present
in the section. The first load level was taken prior to cracking with Vf equal to 76kN (Pf equal
to 152 kN), see Figure 4-25. The concrete stress distribution is approximately quadratic as
expected for plain sections, where the GFRP is not under any significant stress.
78
Figure 4-25 - Shear Stress vs Height (152 kN Pre-Cracking)
Upon cracking, the section no longer behaves as a plane section. The GFRP stirrups begin to
take on the shear load, specifically at the mid-section height of 400 mm and at a height of 112
mm (Figure 4-26). This location corresponds to the beginning of the bend for the stirrups in
the beam. The shear stress at mid-depth is 159 MPa, while at the bend location it is 127 MPa.
Figure 4-26 - Shear Stress vs Height (450 kN SLS)
Upon failure, the shear stress reaches its maximum slightly above the bend location, see Figure
4-27. The bend location is critical due to the reduced ultimate strength and high shear demand.
This is corroborated by the “Load vs. Mid-Span Deflection” response, Figure 4-22, where the
behaviour is uniform at early stages, and only deviates as the section approaches ultimate.
0
100
200
300
400
500
600
700
-0.6 -0.4 -0.2 0 0.2 0.4 0.6
Hei
ght
(mm
)
Stress (MPa)
Concrete Shear Stress
Transverse ReinforcementTensile Stress
0
100
200
300
400
500
600
700
-10 40 90 140 190
Hei
ght
(mm
)
Stress (MPa)
Concrete Shear Stress
Transverse Reinforcement Tensile Stress
79
Figure 4-27 - Shear Stress vs Height (902 kN prior to failure)
JSC–100–50
The analytical and experimental load deformation plots for beam JSC-100-50B, which
contained 100 MPa concrete and 0.5% shear reinforcement using ComBAR stirrups, are
provided in Figure 4-28. Overall the test capacity was captured adequately in the analysis.
However, the deformation was not accurately predicted by VecTor 2. Comparing the two FEM
models, the response was similar at 450 kN, the approximate service load. The difference
between the two analytical curves begins to show at failure, where the bend region appears to
be critical. This is corroborated by the actual mode of failure of the beam, where the bend
region ruptured, see Figure 4-29.
0
100
200
300
400
500
600
700
-10 90 190 290 390 490
Hei
ght
(mm
)
Stress (MPa)
Concrete Shear Stress
Transverse Reinforcement Tensile Stress
80
Figure 4-28 - Load vs Midspan Deflection, VecTor 2 models and experiment
Figure 4-29 – Bend rupture in stirrup (JSC-100-50) (Johnson and Sheikh, 2014)
The reinforcement tensile stress output from Augustus has been provided in Figure 4-30. The
mid depth stress is 452 MPa, equating to a strain of 0.0078 compared to the measured strain of
0.0055. Both these strains are above the ACI limit. The bend stress from Augustus is 254 MPa,
equivalent to a strain of 0.0044.
0
200
400
600
800
1000
1200
1400
0 10 20 30 40 50 60 70
Load
(kN
)
Midspan Deflection (mm)
Experiment
VecTor 2 (Bend)
VecTor 2
81
Figure 4-30 - Augustus shear reinforcement stress at 450 kN (SLS)
Similar to JSV40-40B, beam JSC100-50B shows a quadratic distribution of shear stress for
plain concrete. However, as the section cracks, majority of the stress is carried by the stirrup
near the bend location alone, unlike that of regular strength concrete, see Figure 4-31.
Figure 4-31 - Shear stress along height (450 kN SLS)
0
100
200
300
400
500
600
700
-20 0 20 40 60 80 100 120 140
Hei
ght
(mm
)
Stress (MPa)
Concrete Shear Stress
Transverse Reinforcement TensileStress
82
At failure, it can be seen the bend region is the location of failure where a stress of 531 MPa is
reached. The mid-section, 411 MPa, is only 45% utilized (Figure 4-32). This corroborates the
failure location observed in the experiment.
Figure 4-32 - Shear Stress along Height (1002 kN Failure)
4.4 Modulus Degradation and Creep Strain
It has been shown in some of the referenced literature that a composite fibrous material under
sustained load would see continuous constant creep deformation. The theory, as discussed in
Chapter 2, argues that local fibre failures continue to occur until a global sectional failure takes
place. This phenomenon is important to characterize for two main reasons. Firstly, the different
failure mechanisms, fibre rupture, matrix cracking, and diffusion failure govern sustained load
limits for design. The design should ensure that the sustained load placed on the structure would
not cause a matrix cracking failure during the service life of the structure. Alternatively, if
alkali or water ingress is a concern, maintain load levels well below the diffusion failure level.
Secondly, the reduction of moduli can cause excessive long term deflections causing damage
to partitions and public concern. The following sections provide analytical investigation for the
two topics discussed above and aim to give designers more insight into the behaviour of GFRP
with respects to serviceability limit state design.
0
100
200
300
400
500
600
700
-20 80 180 280 380 480 580
Hei
ght
(mm
)
Stress (MPa)
Concrete Shear Stress
Transverse Reinforcement TensileStress
83
4.4.1 Creep Strain
This section presents representative strain curves from all three test series. Typical figures show
the load history (right Y axis) along with strain data from up to eight gauges (left Y axis), over
the duration of the test (X axis). During the course of the test, some strain gauges began to
malfunction, either stopped providing readings or provided erratic readings. These readings
were removed from the analysis. A comprehensive catalogue of tests is provided in Appendix
A for the reader’s reference.
In order to calculate the creep strain, a definition was required to make readings consistent.
The onset value for creep the creep measurement was defined at the location when the sustained
load was reached, see Figure 4-33 for illustration.
Figure 4-33 - Creep strain calculation illustration
Short Term Tests
Specimens Pt 2 and S 7, Figure 4-34 and 4-35, had test durations of 26 and 10 minutes
respectively, categorizing them as short term tests. Neither specimen showed signs of creep
strain and underwent sudden failures. This contradicts the findings by Mario (2011), whose
tests showed creep strain even at the shortest test durations.
During testing audible ‘pinging’, similar to that of a plucked string, was observed. This was
caused by the individual fibres rupturing. A possible explanation for the lack of creep strain is
84
that creep is only a function of matrix cracking which causes a softer response and leads to
measurable strain over a significant period of time. Failures within this duration and load level
are likely fibre controlled.
Figure 4-34 - Pt2 Strain and Applied Load vs. Time
Figure 4-35 - S7 Strain and Applied Load vs. Time
Mid Term Tests
Specimen S10 (66% of URS) underwent a creep strain of 1400 um/m as seen from Gauge 6 as
shown in Figure 4-36. Load and strain variations with time are shown in Figure 4-37 for
specimen S1 (tested under sustained load of 79% of URS). Strain variation in gauge 6 is drawn
to a magnified class in Figure 4-38 in which a creep strain can be observed under constant load.
85
The creep strain, εcr, was measured to be 700 um/m. Both specimens had test durations in the
10 hr epoch, although they had a difference of 13% in applied load. This should have cause a
difference in TTF of 100 hours. Therefore, it may be that TTF is more directly linked to creep
deformation than applied load. Specimen S1 and S10 both fall under the failure category of
matrix cracking; they were also the only specimens of the S Series to show signs of continued
deformation.
Figure 4-36 - Specimen S10 Strain and Applied Load vs. Time
Figure 4-37 - Specimen S1 Strain and Applied Load vs. Time
86
Figure 4-38 - Specimen S1 Gauge 6 and Applied Load vs Time
The Pt Series had two specimens that fell within the matrix cracking regime, Pt 7 (Figure 4-
39) and Pt 10 (4-40), with a TTF of 44 hours and 52 hours, respectively. Pt 7 and Pt 10 had
measured creep strains (εcr) of 650 and 2000 um/m, respectively.
Figure 4-39 - Pt 7 Strain and Applied Load vs. Time
87
Figure 4-40 - Pt 10 Strain and Applied Load vs. Time
The P Series had three specimens that underwent significant continuous strain, and were
therefore classified under the matrix cracking domain. Specimens P1, P7 and P5 had time to
failures of 8300, 10300 and 12500 seconds. With creep strain values of 2000, 1500 and 1300
um/m all from Gauge 4. See Figure 4-41 for the strain and load history of Specimen P5.
Figure 4-41 - P5 Strain and Applied Load vs. Time
88
Long Term Tests
Specimens within this domain were characterized as undergoing little to no creep deformation
and having time to failure values longer than 1 million seconds or 10 days.
From the S Series, Specimens S8, Figure 4-42, had test duration of 46 days. Unfortunately, due
to pump failures, the loading for S8 was not able to be maintained consistently throughout the
life of the test. A solenoid failure resulted in a pressure spike at the 3.6 million second mark.
However, during the time span from 2 to 3.6 million seconds, it is evident there is no creep
deformation recorded in any of the eight gauges.
Figure 4-42 - S8 Strain and Applied Load vs. Time
Specimen S9, where the loading was maintained fairly uniform over the course of the test
duration of 11 days, did show signs of varying strain measurements (Figure 4-43). However,
if we isolate Gauge 4 and Gauge 6, as shown in Figure 4-44, we can see the former increasing
with the latter decreasing. This could be caused by the concrete creeping and the location of
applied stress shifting away from Gauge 6 to Gauge 4. As a whole there does not appear to be
any clear sign of continuous creep deformation. .
89
Figure 4-43 - S9 Strain and Applied Load vs. Time
Figure 4-44 - Specimen S9 Gauge 4 and 6 vs. Time
Specimen P8 in Figure 4-45, with a TTF of approximately 33 days, had an εcr equal to 580
um/m. Majority of other long term tests showed little to no creep deformation. This may be
due to the specimens being in the diffusion failure domain where micro cracks coalesce
eventually merging would cause failure.
90
Figure 4-45 - Specimen P8 Strain and Applied Load vs. Time
A summary of all specimens, with creep strain values with accompanying gauge number and
time to failure is presented in Table 4-11. It can be seen that specimens which failed within
1000 seconds or beyond 1 million seconds showed no creep deformation. These two time
values are the approximate boundaries for the three regions of creep rupture failure.
91
Table 4-11 - Creep strain summary of sustained load specimens
When all creep strain values are plotted against their respective TTF values, it can be seen that
three “zones” or regions exist corresponding to the 10,000-second and 1,000,000-second
domains. All specimens were plotted together as to show a general trend (Figure 4-46). There
are simply not enough test specimens which span every time epoch to be able to know with
certainty when one region ends and another begins.
Specimen
Creep
Strain
(um/m) Gauge
TTF
(sec)
P1 2000 G7 8300
P7 1500 G6 10350
P5 1200 G6 12500
P6 0 N/A 2210000
P8 800 G6 2850000
P9 0 N/A 4700000
P2 0 N/A 9100000
Pt2 0 N/A 1163
Pt6 0 N/A 157000
Pt7 650 G6 159000
Pt10 2000 G4 186500
Pt1 800 G4 1147000
Pt8 0 G4 1275000
S7 0 N/A 580
S2 0 N/A 829
S10 1400 G6 30000
S1 700 G6 41600
S9 0 N/A 964000
S6 0 N/A 1631000
S8 0 N/A 3710000
92
Figure 4-46 - Total Creep Strain vs. Time to Failure
The regression equations in Table 4-7 were isolated for the variable x (sustained stress). The
applied load associated with 10,000-seconds and 1,000,000-seconds was calculated for all three
series. From this, the average % of URS values were calculated to create Figure 4-47. The
regions, red, yellow and green, correspond to the failure domains of fibre rupture, matrix
cracking, and diffusion failure, respectively. The corresponding stress values for the borders
for the three regions are 0.77 and 0.65 of URS. This is to serve as a gross approximation of
where the three regions exist for the type of GFRP Stirrups tested. However, individual regions
would vary for each type of material. This corroborates the findings of Sa et al. (2011), who
showed specimens loaded to 60 to 80% of URS showed creep and specimens loaded to 50%
of URS and lower did not.
0
500
1000
1500
2000
2500
10 100 1000 10000 100000 1000000 10000000
Cre
ep S
trai
n (
um
/m)
Time to Failure (sec)
93
Figure 4-47 - Creep rupture domains
4.4.2 Moduli Degredation
To evaluate the degradation in the modulus of elasticity, the applied load was cycled a few
times during the period of sustained loading. The original intent was to apply the sustained load
for a duration equal to approximately 75% of the projected TTF, at which point the load was
removed and reapplied over a course of several minutes. The initial modulus was compared to
the modulus during reloading. Due to the variability of the time to failure (TTF), some
specimens failed prior to having their applied load cycled. Conversely, some specimens failed
much later than expected and the applied load was cycled much earlier in their lifespan. In this
section, several specimens and their respective “Stress vs. Strain” figures are discussed. A
comprehensive list of all tests can be found in the Appendix. Figure 4-48 illustrates the stress-
strain behavior obtained from the load cycles.
94
Figure 4-48 - Theoretical stress vs. strain behaviour
Specimen Pt8 had a TTF of 354 hours with the load cycle occurring after 104 hours of sustained
loading (460 MPa). From Figure 4-49 it is evident that there was no creep strain. The strain
recovery in this specimen (Figure 4-48) was approximately 800 um/m. The initial modulus was
measured to be approximately 41.2 GPa, with the secondary modulus of 46.8 GPa. The
increased modulus after a period of sustained load seems to be an anomaly.
95
Figure 4-49 - Pt8 - Stress vs. Strain (Gauge 3)
A sustained load of 512 MPa was applied for approximately 700 seconds in specimens Pt2,
after which the specimen was unloaded and reloaded. Upon the reaching the sustained load,
failure occurred shortly after (1160 sec.). The initial and secondary modulus were both
measured to be 46.8 GPa (Figure 4-50).
Figure 4-50 - Pt2 - Stress vs. Strain (Gauge 3)
Specimen P5, loaded under a sustained stress of 286 MPa, measured an initial modulus of 41.3
GPa, and creep strain of approximately 1200 um/m after 3.22 hours of sustained loading
(Figure 4-51). Upon unloading, the modulus followed a slope similar to that of the initial
Unload/Reload - 104 h
Failure - 354 h
0
100
200
300
400
500
600
0 2000 4000 6000 8000 10000 12000
Stre
ss (
MPa
)
Strain (um/m)
Gauge 3
Failure - 1160 sec
Unload/Reload - 750 sec0
100
200
300
400
500
600
0 2000 4000 6000 8000 10000 12000
Stre
ss (
MPa
)
Strain (um/m)
Gauge 3
96
loading (41.3 GPa). However, the secondary modulus was measured to be 40.3 GPa; failure
followed shortly after 0.25 hours of additional sustained loading.
Figure 4-51 - P5 - Stress vs. Strain (Gauge 3)
Specimen P6, loaded under a sustained stress of 263, had an initial stress-strain relationship
(Figure 4-52) which did not follow a linear trend, however, the initial tangential modulus was
measured to be 50.7 GPa. An accurate measurement of creep strain was not possible due to
difficulties with consistent strain readings from the gauge. After 472 hours of sustained
loading, the secondary modulus was measured to be approximately 40.0 GPa.
Unload/Reload - 3.22 h
Failure - 3.47 h
0
50
100
150
200
250
300
350
0 1000 2000 3000 4000 5000 6000 7000 8000 9000
Stre
ss (
MPa
)
Strain (um/m)
Gauge 3
97
Figure 4-52 - P6 - Stress vs. Strain (Gauge 3)
Specimen P9 had a TTF of 578 hours with the load cycle occurring after 534 hours of sustained
loading (267 MPa) (Figure 4-53). No significant creep strain was apparent, however, upon
unloading, there was a strain recovery of approximately 500 um/m. The initial modulus was
measured to be approximately 77.0 GPa, with a similar secondary modulus of 77.5 GPa.
Figure 4-53 - P9 - Stress vs. Strain (Gauge 3)
Specimen S1, loaded under a sustained stress of 528 MPa, had an initial modulus at the
location of strain gauge 6 was measured to be 31.8 GPa (Figure 4-54), which is significantly
lower than the measured moduli for other specimens presented partly due to the fact that this
Unload/Reload - 472 h
Failure - 584 h
0
100
200
300
400
500
0 2000 4000 6000 8000 10000 12000 14000 16000
Stre
ss (
MPa
)
Strain (um/m)
Gauge 3
Unload/Reload - 534 h
Failure - 578 h
0
100
200
300
400
500
600
0 1000 2000 3000 4000 5000 6000 7000 8000
Stre
ss (
MPa
)
Strain (um/m)
Gauge 3
98
gauge is located right after the bend. The low modulus may also be an anomaly in the strain
readings, however, our focus is to obtain a relative sense of the initial and secondary
modulus. The measured creep strain was approximately 700 um/m. The specimens failed at
11.6 hours and the location of failure was near gauge 7. The modulus recorded during the
final unloading at gauge 6 location was slightly lower than the initial 30.8 GPa.
Figure 4-54 - S1 - Stress vs. Strain (Gauge 6)
S2 was loaded to a sustained stress of 503 MPa. The initial modulus calculated from gauge 4
for S2 (Figure 4-55) was 35.7 GPa. The slopes of the unloading and reloading curves were
similar and the calculated modulus was 34.2 GPa with an approximate creep strain of 2000
um/m and creep strain recovery of 500 um/m. Note that the creep strain measurement from
gauge 4 was not shown in Table – 11 as the strain readings were intermittent during the
sustained loading portion of the test.
Failure - 11.6 h
0
100
200
300
400
500
600
0 2000 4000 6000 8000 10000 12000 14000 16000 18000 20000
Stre
ss (
MPa
)
Strain (um/m)
Gauge 6
99
Figure 4-55 - S2 - Stress vs. Strain (Gauge 4)
Specimen S8 was subjected to variable loading due to a pressure switch failure after
approximately 888 hours of sustained loading (427 MPa), see Figure 4-56. No cyclic loading
was conducted on the specimen. The initial modulus and the modulus measured during
unloading, after failure (1031 h), were calculated to be 58.0 GPa in both instances. During the
load increase at elapsed time of 888 hours, the stress-strain slope was much steeper most likely
due to an erroneous reading of the strain gauge and did not follow an expected trend.
Figure 4-56 - S8 - Stress vs. Strain (Gauge 7)
Unload/Reload -600 sec
Failure - 830 sec
0
100
200
300
400
500
600
0 2000 4000 6000 8000 10000 12000 14000 16000
Stre
ss (
MPa
)
Strain (um/m)
Gauge 4
Load Increase - 888 h
Failure - 1031 h
0
100
200
300
400
500
600
700
0 2000 4000 6000 8000 10000 12000
Stre
ss (
MPa
)
Strain (um/m)
Gauge 7
100
From the comparison of initial moduli to that of the secondary moduli for specimens presented,
there appears to be no strong pattern of degradation. Aside from specimen P6’s modulus
reduction of 20% all other specimens were able to maintain their initial values or only see a
small reduction of 2-5%.
4.5 Cumulative Weakening Model
The creep rupture regressions provided in Table 4-7 are based on the average load values and
do not take into account the exponential nature of damage caused by sustained stresses. An
average value assumes damage caused by an applied load is linear. The CW model is applied
to Specimen S8 to provide an “equivalent time”. This is necessary to capture the varying
damage caused by variable loading. The load history for specimen S8 has been provided in
Figure 4-57. For illustrative purposes it has been divided into three different regions. Region
1, highlighted in green, has an applied load lower than the average value of 103 kN, therefore
the equivalent time is 1.1 million seconds as opposed to the unadjusted time of 1.3 million
seconds. Region 2, highlighted in yellow, has an applied load equal to that of the average load
and remains unaltered, and its equivalent time is equal to the unadjusted time of 1.9 million
seconds. Region 3, highlighted in red, which has an applied load higher than the average has
an equivalent time of 6 million seconds as opposed to the unadjusted time of 800,000 seconds.
For the actual calculation of equivalent time, 1 second time steps were taken, therefore
specimen S8 was numerically integrated over 3.9 million steps.
Figure 4-57 - Equivalent time illustration for S8
101
The Cumulative Weakening Model as developed by Franke and Meyer (1992) was modified
to work for the three series of stirrups tested. In order to find the equivalent time Equation 2-
4 was used with the values for specimen S8 inserted. The value of ns was taken as the creep
stress compliance (temporal slope) value as summarized in Table 4-2 for the three series.
𝑡1 =1
𝜎1𝑛𝑠 ∫ 𝜎(𝑡)𝑛𝑠𝑑𝑡
𝑡𝑏0
[2-4]
𝑇𝐶𝑊 =1
554.8920.59∫ 𝜎(𝑡)20.59𝑑𝑡3.9𝐸6
0
In order to calculate the cumulative weakness, the value of TCW was divided by the actual TTF
(TA). A value of TCW/TA below one implies a specimen failed prematurely, a value above 1
implies a specimen possessed higher capacity than expected. Due to the variability in the
strength of GFRP bent bars tested and the magnification of the exponential value of ns, the
cumulative damage values varied greatly. Additionally, the slope of the creep rupture curve is
extremely shallow. The slightest change in loading yields failures in a completely different
time epoch. Perhaps if the model were to be applied to the straight bar tests conducted by Weber
(2005), the cumulative weakening model would be of additional value. Not only were the test
results less varied, R2 value of 0.93, but they also possessed a higher creep stress compliance
(ns), resulting in a more prominent correlation between applied load and time to failure.
The cumulative weakening model was applied to all 18 specimens; only four specimens had
varied loading due to equipment issues and the remaining 14 specimens had fairly constant
loading throughout the loading period. The application of the CWM to the consistent 14 tests
was done to calibrate the model and to check the validity of the theory. Specimens S8, P1, P2
and Pt8 possessed some variation in loading, and once the model was applied, the equivalent
time was found to be greater than that of the unadjusted time, signifying that the damage
sustained by the specimens was larger than the simple linear regressions. The CWM adjusted
TTF values have been graphed with the experimental results in Figures 4-58, 4-59 and 4-60.
Note that in these figures, there are equal number of points from the experimental and the
analytical results but they may be overlapped. The differences in the P and Pt Series were found
to be insignificant however, it shows the S Series to be more resilient than previously calculated
102
as the variable load had dealt more damage to S8 than was accounted for with the simple
average load calculation.
Figure 4-58 - S Series Applied Stress vs. Time to Failure with cumulative weakening model
Figure 4-59 - P Series Applied Stress vs. Time to Failure with cumulative weakening model
0
100
200
300
400
500
600
700
800
1.E-02 1.E-01 1.E+00 1.E+01 1.E+02 1.E+03 1.E+04
Ap
plie
d S
tres
s (M
Pa)
Time to Failure (hr)
CWM
Experimental
0
100
200
300
400
500
1.E-04 1.E-03 1.E-02 1.E-01 1.E+00 1.E+01 1.E+02 1.E+03 1.E+04
Ap
plie
d S
tres
s (M
Pa)
Time to Failure (hr)
CWM
Experimental
103
Figure 4-60 - Pt Series Applied Stress vs. Time to Failure with cumulative weakening model
A comparison of cumulative weakening values to TA/TP values have been provided in Figure
4-61. Specimens where both values were equal, the loading was fairly constant. Take Pt1, for
example, the unadjusted and equivalent time are nearly identical, 1,143,000 and 1,146,000
seconds respectively. Where TA/TP and TA/TCW value varied, the specimen, such as S8, had
variable loading.
Figure 4-61 - Cumulative weakening comparison to TA/TP
Since the TA/TP values do not take into account the logarithmic nature of the TTF curve of FRP
specimens under sustained loading, it is less accurate than that of the CWM model. With the
0
100
200
300
400
500
600
700
1.E-02 1.E-01 1.E+00 1.E+01 1.E+02 1.E+03
Ap
plie
d S
tres
s (M
Pa)
Time to Failure (hr)
CWM
Experimental
104
CWM model calibrated it could potentially be used to model FRP specimens with highly
variable loading such as fatigue or cyclic loading.
105
5 Conclusion
Thirty stirrup specimens from three manufacturers were tested under instant load and constant
sustained loads which were fractions of the instant ultimate capacity. In four of the 30
specimens, the sustained load varied slightly which was accounted for in the cumulative
weakening model analysis. Specimens were monitored constantly until failure. The sustained
loads varied between 66% and 93% of their respective ultimate capacities. One of the main
goals was to ascertain the stress levels for 100-year service life of these stirrups. A number of
conclusions can be made from this study and are discussed below.
The JSCE equation, 2-5, for calculating bend strength predicted the bend strength for the Pt
stirrups reasonably accurately (99%). The predictions for P and S stirrups were accurate to
about 75% 142%. The bars tested in this study were typical of the industry. However, it can be
concluded that there is substantial variability in the bent bar strengths which can present a
challenge.
From the data on a limited number of tests, the stress levels for 100-year service life varied
between 40% and 49% of the ultimate strength of stirrups. In absolute terms the stress levels
varied between 193 MPa and 270 MPa.
The strain limits for sustained loading imposed by ACI 440 and CSA S806 appear to be
adequately conservative for projected strengths. However, once the variability of GFRP
stirrups is taken into consideration there is a concern with potential premature failure. This
issue is further exasperated when coupled with high shear stresses at the bend location, even at
SLS conditions. The strain limits imposed by ACI 440 and CHBDC S6 are to preserve
aggregate interlock for ultimate limit states and not specifically creep rupture.
The beams tested by Johnson (2014) were designed to be shear critical and are not typical of
standard beams which would be designed by an engineer. Beams are typically designed to have
low shear utilization. However, the beam with the shear reinforcement ratio of 0.4% and stirrup
spacing of 150 mm was found to be susceptible to premature shear failure. The shear
reinforcement in this beam is fairly significant and represents the lower range of stirrup
spacing. The conclusions made from the analysis conducted are only valid for beams of this
nature, stirrup dimensions, GFRP types, and loading conditions.
106
From the findings in this study, and the literature review provided, it can be resolved that
continued creep deformation is not of concern especially at the lower levels of sustained stress.
Furthermore, the GFRP materials tested do not display moduli degradation. The issue of creep
strain in GFRP appears to only affect the older generation which used polyester resins and not
the current generation of vinyl ester resins. Any long term deflection to be seen in a structure
would most likely be caused by concrete undergoing creep.
The cumulative weakening model developed by Franke and Meyer (1992) was calibrated for
the three series of GFRP stirrups. From this model, it is possible to account for variable loading
and could be used in applications such as fatigue loading.
107
6 Recommendations for Future Work
There are several areas in which GFRP research should be advanced for sake of designing safer
structures and for making GFRP a more attractive alternative.
Long term testing of beams with high strength concrete to determine if the findings in this
study are valid.
Generate test data from statistically large sets of bent bars in order to better understand the
distribution of ultimate strength. This would aid in the development of serviceability strain
limits based on the probable premature failure in addition to aggregate interlock which is the
current motivation behind the ACI 440 limit.
With the current design practice GFRP does not provide a significant incentive over traditional
steel reinforcement for concrete construction. One area, in which FRP reinforcement could
have a significant advantage is parking garages. Currently, parking garages require a very
expensive waterproofing layer on suspended slabs to prevent corrosion of steel reinforcement.
If it can be conclusively shown that a waterproofing membrane is not required for FRP, it
would provide a large upfront cost savings to the owner. The studies would need to look at the
effect of ingress of hydrocarbons, and water. Furthermore, the testing would need to look at
the crack widths to prevent leaking of water through suspended slabs and potential damage
caused by freeze thaw.
The specimens in this study underwent not only tensile loading but also the inherent shear
loading at the bend location. The combined effect of tensile and shear loading should be
analysed on FRP materials. This is especially important due to the anisotropic behaviour of
FRP materials and its weakness in the transverse direction.
108
7 References
ACI 440.1R-015, 2015 “Guide for the Design and Construction of Structural Concrete
Reinforced with FRP Bars”, ACI Committee 440. American Concrete Institute, Farmington
Hills, Mich. U.S.A.
Almusallam T. and Al-Salloum Y., 2004, “Durability of GFRP Rebars in Concrete Beams
under Sustained Loads at Severe Environments”, Journal of Composite Materials, Department
of Civil Engineering, King Saud University, Saudi Arabia.
ASTM D2990-09, 2009 “Standard Test Methods for Tensile, Compressive, and Flexural Creep
and Creep Rupture of Plastics.”, American Standards and Testing of Materials. 100 Barr
Harbor Drive, PO Box C700, West Conshohocken, PA 19428-2959. United States
Bagonluri-Nuuri, David Fred. 1998, “Simulation of Fatigue Performance & Creep
Rupture of Glass-Reinforced Polymeric Composites for Infrastructure Applications”. Master
of Engineering Thesis. Faculty of the Virginia Polytechnic Institute and State University.
Bank LC, Mosallam AS. 1992, “Creep and failure of a full-size fibre reinforced plastic
pultruded frame.” Compos Eng. 2(3):213–27.
Bank LC, 2006 “Composites for Construction: Structural Design with FRP Materials.”
Hoboken, NJ: John Wiley & Sons. Print.
Canadian Standards Association, 2014 “Canadian Highway Bridge Code”, CSA S6-14,
Mississauga, Canada.
Canadian Standards Association, 2012, Design and Construction of Building Components with
Fibre Reinforced Polymers. CSA S806-12, CSA Mississauga, Canada.
Canadian Standards Association, 2012 “Specifications for Product Certification of FRP as
Internal reinforcements in Concrete Structures, CSA S807-10, Mississauga, Canada.
Canadian Standards Association, 2012 “Specification for Fibre-Reinforced Polymer (FRP)
Materials for Externally Reinforcing Structures, CSA S808-14, Mississauga, Canada.
Daniali S., “Short-term and long-term behavior of two types of reinforced plastic beams.” 46th
Annual Conference, Composites Institute; February1991:13-A1-5.
Franke L., Meyer J., 1992, “Predicting the Tensile Strength and Creep-Rupture Behaviour of
Glass-Reinforced Polymer Rods”, Journal of Material Science, Technical University of
Hamburg-Harburg, Hamburg, Germany.
Hays, G.F., 2010. NACE-International, The corrosion society.
http://events.nace.org/euro/corrodia/Fall_2010/wco.asp
ISIS Canada, 2007, “Reinforcing Concrete Structures with Fibre Reinforced Polymers”. ISIS
Canada Design Manual, University of Manitoba, Winnipeg.
Japan Society of Civil Engineers (JSCE), 1997b, “Recommendation for Design and
Construction of Concrete Structures Using Continuous Fiber Reinforcing Materials,” Concrete
Engineering Series No. 23, 325 pp.
109
Johnson, D.T, 2009. "Investigation of Glass Fibre Reinforced Polymer Reinforcing bars as
Internal Reinforcement for Concrete Structures", Master's Thesis, Department of Civil
Engineering, University of Toronto.
Johnson, D. T. C., 2014, “Investigation of Glass Fibre Reinforced Polymer (GFRP) Bars as
Internal Reinforcement for Concrete Structures”, Ph.D. Thesis, Department of Civil
Engineering, University of Toronto, Toronto, Canada.
Johnson, D. T. C. and Sheikh, S. A. 2013. Performance of bent stirrups and headed bars in
concrete structures. Canadian Journal of Civil Engineering, Oct. 2013, Vol. 40, No. 1, pp.
1082-1090.
Jun Koyanagi , Genya Kiyota , Takashi Kamiya & Hiroyuki Kawada (2004):
Prediction of creep rupture in unidirectional composite: Creep rupture model with interfacial
debonding and its propagation, Advanced Composite Materials, 13:3-4, 199-213
Laoubi K., El-Salakawy E., and Benmokrane B., “Creep and durability of sand-coated glass
FRP bars in concrete elements under freeze/thaw cycling and sustained loads”, Cement and
Concrete Composites, September 12th, 2006, Department of Civil Engineering, University of
Sherbrooke, Sherbrooke, Quebec, Canada.
McBagonluri-Nuuri, D. F., 1998, “Simulation of Fatigue Performance and Creep-Rupture of
Glass-Reinforced Polymeric Composites for Infrastructure Applications, MS Thesis, Virginia
Polytechnic Institute and State University, Blacksburg, Virginia.
Mukherjee A. and Arwikar SJ. “Performance of glass fiber reinforced polymer reinforcing bars
in tropical environments-Part I: Structural scale tests”. ACI Structural J 2005; 102(5): 745–
753.
Nkurunziza G. and Benmokrane B. (2005), “Effect of Sustained Load and Environment on
Long-Term Tensile Properties of Glass Fibre-Reinforced Polymer Reinforcing Bars”, ACI
Structural Journal, Technical Paper, Vol. 102, No 4., July-August 2005.
Sheikh, S. A. and Johnson, D. T. C., “Performance of Bent Stirrup and Headed GFRP Bars in
Concrete Structures,” The 6th International Conference on Advanced Composite Materials in
Bridges and Structures (ACMBS-VI), May 22-25, 2012, Kingston, Ontario.
Weber, A., 2007, “16mm Tensile Test Report 25.05.2007”, Report No: 075 07, Schöck
Research Labs, Baden Baden, Germany.
Weber, A., 2007, “8mm Tensile Test Report 25.05.2007”, Report No: 075 07 ComBAR-8,
Schöck Research Labs, Baden Baden, Germany.
Weber, A., 2007, “12mm Tensile Test Report 25.05.2007”, Report No: 075 07 ComBAR12,
Schöck Research Labs, Baden Baden, Germany.
Weber, A., 2007, “Expert Report, Application for DIBt Certification of GFRP Reinforcing
Bars “ComBAR” made by Schöck Bauteile GmbH”, Schöck Research Labs, Baden Baden,
Germany.
1
Appendix A - Test Results P Series Strain and Applied Load versus Time
2
3
4
5
Pt Series Strain and Applied Load versus Time
6
7
8
9
10
S Series Strain and Applied Load versus Time
11
12
13
14
15
S Series - Strain vs Stress
16
17
18
P Series – Stress vs. Strain
0
50
100
150
200
250
300
0 1000 2000 3000 4000 5000 6000 7000 8000 9000
Stre
ss (
MPa
)
Strain (um/m)
P5 - Stress vs. Strain
Gauge 3
19
0
50
100
150
200
250
300
350
0 2000 4000 6000 8000 10000 12000 14000
Stre
ss (
MPa
)
Strain (um/m)
P7 - Stress vs. Strain
Gauge 3
20
0
50
100
150
200
250
300
0 2000 4000 6000 8000 10000 12000 14000
Stre
ss (
MPa
)
Strain (um/m)
P8 - Stress vs. Strain
Gauge 6
21
Pt Series - Strain vs Stress
22
23
0
100
200
300
400
500
0 2000 4000 6000 8000 10000 12000 14000
Stre
ss (
MPa
)
Strain (um/m)
Pt10 - Stress vs. Strain
Gauge 6
24
Appendix B – Measurement Equipment Information
25
26
Appendix C - Strain Gauge - TML UFLA 2-350-11