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Shear Connection Systems and Interfacial Stresses
in GFRP
Lu
M.Sc. Thesis Extended Abstract
Shear Connection Systems and Interfacial Stresses
in GFRP-Concrete Hybrid Beams
Luís Daniel Tavares Nogueira
M.Sc. Thesis Extended Abstract
November 2009
Shear Connection Systems and Interfacial Stresses
Concrete Hybrid Beams
1
1 Introduction
A composite material can be defined as a material which results from the combination of two or
more materials and maintains an identifiable interface surface. Fiber reinforced polymers
(FRPs) are amongst these composite materials.
FRPs have been used in the construction industry since the Second World War, although it was
only in the second half of the 1990´s that their application in civil engineering grew significantly
[1]. Nowadays, the use of FRPs in construction is quite generalized, and they are used in both
structural and non structural applications. In the structural field, FRPs have been applied in the
construction of new structures substituting steel, whether as a concrete reinforcement or as the
primary structure material [2-4]. The use of FRPs has also grown in the rehabilitation sector
where, frequently, it allows maintaining the historical context while adapting a structure to
nowadays serviceability and safety requirements [5].
The most used FRPs in the construction industry are glass fiber reinforced polymers (GFRPs).
Their use is justified by the high strength-weight ratio, the electromagnetic transparency, the
chemical resistance and the relatively low production costs. However, in addition to a low
resistance when submitted to high temperatures, their low elasticity modulus makes them
relatively deformable and susceptible to instability phenomena. In order to prevent the
occurrence of instability problems and to improve the stiffness of composite structural elements,
GFRP can be combined with traditional materials, such as concrete and steel, creating hybrid
(or composite) structures with a reasonable strength-weight ratio and stiffness.
2 GFRP-concrete hybrid beams
Combining two distinct materials in order to take advantage of their best properties is a
technique which has been for used for ages. In the past two decades many studies have been
conducted in order to find the best way to combine the compressive properties of concrete and
the tensile resistance of GFRP [6-13]. This combination is, however, hard to materialize due to
the GFRP´s low elasticity modulus, which makes it prone to instability phenomenon and due to
the debonding of the soffit reinforcement when submitted to bending [14]. Debonding occurs
when a horizontal crack develops in the beam´s end and then spreads towards the midspan
section, causing the separation of the two materials. This cracking occurs because, when the
composite beam is submitted to bending, the concrete deforms, while the soffit reinforcement
tries to maintain its initial position, therefore creating a stress peak at the extremity of the beam.
Practically, this can occur either by concrete cover separation, if the cracking develops along
the rebar plane, or by plate end interface debonding, if the cracking occurs leaving a thin layer
of concrete on the plate, with the latter mechanism being far more uncommon [15].
The objective of this work wa
beams when submitted to bending and
cracking from spreading along the beam
used in [13] and it consists
(200×100×10 mm). The beam´s geometry is
system two rows of M10 bolts placed near the beam´s end
Figure 1 - Proposed solution
3 Evaluation of interfacial s
Debonding failure may occur when a peak of both normal and shear stresses arises
extremity of the beam. Therefore, these stresses distributions and
in order to understand the debonding
models (FEM) were built using Ansys Multiphysics
experimentally tested by Correia [13] and
Error! Reference source not found.
therefore, in each case, only half of the beam was modeled. Eight node prismatic elements
were used and the mesh was
such as the load application area, near the supports a
latter location demands a rather dense mesh due to the singularity created by the concrete
adhesive interface [15,16]. A 0.25
Figure 3 - General view of the FEM model.
was to study the debonding mechanism in GFRP
when submitted to bending and to develop an anchoring system able to
along the beam. The structural shape adopted is the same as the one
ts in a 10×40 cm2 concrete slab bonded to a GFRP I
The beam´s geometry is represented in Figures 1 and 2. For the anchor
system two rows of M10 bolts placed near the beam´s end were adopted.
Proposed solution. Figure 2 - Lateral view of the proposed solution
of interfacial stresses
occur when a peak of both normal and shear stresses arises
. Therefore, these stresses distributions and their values are crucial data
in order to understand the debonding mechanism. To assess their values, five 3D finite element
were built using Ansys Multiphysics software. The modeled beams were the
tested by Correia [13] and their geometries and test results are summarized in
Error! Reference source not found.. Symmetry considerations have been taken into account,
therefore, in each case, only half of the beam was modeled. Eight node prismatic elements
were used and the mesh was refined in regions were stress concentrations were expected,
such as the load application area, near the supports and at the extremity of the
latter location demands a rather dense mesh due to the singularity created by the concrete
16]. A 0.25 mm mesh was therefore adopted at the end of the
General view of the FEM model. Figure 4 - Suport conditions of the FEM model.
2
GFRP–concrete hybrid
able to prevent the
. The structural shape adopted is the same as the one
to a GFRP I beam
For the anchoring
Lateral view of the proposed solution.
occur when a peak of both normal and shear stresses arises at the
values are crucial data
five 3D finite element
. The modeled beams were the ones
are summarized in
een taken into account,
therefore, in each case, only half of the beam was modeled. Eight node prismatic elements
were stress concentrations were expected,
extremity of the beam. This
latter location demands a rather dense mesh due to the singularity created by the concrete-
end of the beam.
Suport conditions of the FEM model.
3
Table 1 - Correia´s flexural texts results summary (adapted from [13]).
Beam Support conditions Load arrangements Failure mode
HB2 4.0 m span
simply supported 1 load, midspan Concrete - adhesive interface
HB4a 1.8 m span
simply supported 2 loads, 1/3 span GFRP web shear
HB4b 1.8 m span
simply supported 2 loads, 1/3 span Concrete - adhesive interface
HB5 1.8 m
simply supported 2 loads, 1/3 span GFRP web shear
HB7 two 2.8 m spans 2 loads (at 3/8 each span)
1st
/ 2nd
- Concrete - adhesive
interface
Final – GFRP web crushing
The study of previously tested beams included the comparison of midspan displacements and
both shear and normal stresses distributions along a path through the beam´s longitudinal axis,
at a depth of 0.5 mm inside the concrete slab. This path and depth corresponded roughly to the
development of the longitudinal crack, which caused the tested beams to collapse. The normal
stresses were compared with the concrete average maximum tensile stress (fctm), using three
different reference criteria: maximum normal stress (σy,max), extrapolated normal stress (σy,ext),
which corresponded to the stress distribution in the last 100 mm, not considering the peak at the
end of the beam, and average normal stress (σy,med) which corresponded to the average normal
stress at the beam´s extremity section.
Figure 5 - Extrapolated normal stresses in HB2
beam.
Figure 6 - Average normal stresses at the HB2
beam extremity section.
The numerical analyses using the finite element models, confirm that a normal stress peak
occurs at the end of the beam. This peak, which is caused by the beam´s end singularity, is 3.0
to 4.5 times higher than fctm, except in beam HB4a, which collapsed due to web crushing. It is,
therefore, too conservative to predict the beam´s ultimate load using the maximum normal
stress criterion when the beam is submitted to pure bending.
0,0
2,0
4,0
6,0
8,0
10,0
0255075100
No
rma
l str
ess
es,
σσ σσzz
[MP
a]
x [mm]
Normal
Extrapolation
0,0
2,0
4,0
6,0
8,0
10,0
150 175 200 225 250
No
rma
l str
ess
σσ σσ
yy
[MP
a]
z [mm]
Normal
Average
4
Figure 7 – HB2, HB4b and HB7 normal stresses
near the beam end.
Figure 8 - HB4b and HB5 normal stresses near the
beams end.
Although this analysis conducted to lower normal stresses (compared to those corresponding to
the maximum normal stresses), these stresses were higher than fctm, except in beams HB4a
and HB4b, where the extrapolated stresses were 47.0% and 4.7% lower than fctm, respectively,.
On the other beams, the extrapolated normal stresses were between 56.5% to 135% higher
than fctm. It is therefore possible to conclude that the linear extrapolation of normal stresses in
the beam´s extremity is a conservative method of predicting the ultimate load. This result
becomes evident in beam HB5, which collapsed due to web crushing, even though the
extrapolated stress was 76% higher than fctm.
The average normal stress at the end of the beam was the criterion that provided the closest
results, compared to fctm. This method of analysis was the one which produced the lower normal
stresses, which were still 10% to 50% higher than fctm. As for the extrapolated stresses, the
exceptions were beams HB4a and HB4b, both presenting average normal stresses lower than
fctm. In the former beam, results can be explained by the fact that it collapsed due to web
crushing, not presenting any crack in its extremities. In the latter beam, which collapsed without
any prior visible warning, the applied load corresponded to the occurrence of audible cracks,
presumably at the beam´s interface, resulting, for that reason, in a normal stress lower than fctm.
In what concerns shear stresses, their maximum values take place in the supports vicinity,
which is where the shear load is maximum. Near the extremity, usually less than 10 mm from
the end, shear stress peaks are also observed. However, these stresses are much lower than
those registered near the supports. In the modeled beams, the shear stresses in the extremities
presented values close to zero, which meets the condition of zero shear stresses in a free
surface.
0,0
2,0
4,0
6,0
8,0
10,0
020406080100
No
rma
l st
ress
es,
σσ σσzz
[M
Pa
]
x [mm]
HB2
HB4b
HB7
fctm
-5,0
0,0
5,0
10,0
15,0
20,0
020406080100
No
rma
l st
ress
es,
σσ σσzz
[M
Pa
]
x [mm]
HB4a
fctm HB4a
HB5
fctm HB5
5
Figure 9 - Shear stress distribution near the beam
end in beams HB2, HB4b and HB7.
Figure 10 - Shear stress distribution near the beam
end in beams HB4a, HB5.
The vertical displacements obtained using the FEM were 22% lower than the ones measured by
Correia in the test of beam HB2 [13]. This difference was caused, most likely, due to the fact
that the effect of concrete cracking was not considered in the FEM, which performed linear
elastic analyses. This is shown in beam HB5, where the neutral axis was below the interface
(which means that all concrete slab was under compression) and the midspan displacements
are 10.7% lower than the experimental ones. Other factors that may explain the displacement
discrepancy are the inaccuracy in the estimation of the GFRP and concrete elasticity moduli.
4 Experimental tests
4.1 Materials
The concrete used in the experimental tests had an average compressive strength of 48.7 MPa,
an average tensile strength of 3.3 MPa and an elasticity modulus of 32.8 GPa. For the bonded
connection systems a MBrace Resin 220 epoxy resin was used. This resin has a tensile
strength of 33.0 MPa, a 4.8 x10-3
strain at failure and an elasticity modulus of 7.5 GPa [17]. The
anchoring system was materialized through M10 with a class of resistence of 8.8. The pultruded
GFRP I-profile had the following nominal dimensions: height = 200 mm; width = 100 mm; web
thickness = 10 mm; flange thickness = 10 mm. The mechanical properties of the GFRP profile
were determined in [2,18] through tests performed in similar profiles. Table 2 presents average
values, for the different mechanical properties: ultimate stress (σu), Young’s modulus (E),
ultimate strain (εu), Poisson ratio (νxy) and interlaminar shear strength (Fsbs).
-1,0
-0,5
0,0
0,5
1,0
1,5
2,0
01020304050
Sh
ea
r st
ress
es,
ττ ττx
y [
MP
a]
x [mm]
HB2HB4bHB7
-2,0
-1,0
0,0
1,0
2,0
3,0
4,0
01020304050
Sh
ea
r st
ress
es,
ττ ττx
y [
MP
a]
x [mm]
HB4a
HB5
6
Table 2 - GFRP profile mechanical properties (adapted from [2]).
Flexural Tensile Compression
Longitudinal [19] Longitudinal [20] Longitudinal [21] Transversal [21]
σu [MPa] 624.6 475.5 375.9 122.0
E [GPa] 26.9 32.8 26.4 7.4
εu [x10-3
] 24.9 15.4 17.0 21.5
νxy - 0.28 - -
Fsbs = 35.0 MPa [22]
4.2 Push out tests
Push out tests were carried out in order to assess the behavior of three different shear
connection systems between the GFRP profile and concrete: (i) bolted connections (P)
materialized with two M10 steel bolts on each flange; (ii) bonded connections (C) in which the
GFRP profile was glued to concrete with a 2 mm thick epoxy adhesive layer; (iii) and bonded-
bolted connections (CP) consisting in the combination of the two aforementioned types of
connections. The tested samples consisted of a 34 cm long GFRP profile connected to two
20 cm concrete cubes.
Load was applied at the top of the GFRP profile with a Enerpac hydraulic jack with a load
capacity of 200 kN. The applied load was measured with a load cell from Novatech with a load
capacity of 200 kN, placed between the top of the GFRP profile and the hydraulic jack. The
relative displacements between the profiles and the concrete cubes were measured using an
electrical displacement transducer at each vertex of a metallic plate placed atop the GFRP
profile. For each connection system three samples were tested. The test set-up used in the
push out experiments is represented in Figure 11 and 12.
Figure 11 - Test set up used in the push out tests. Figure 12 – Displacements measurement setup
The tested systems were compared according to the following three criteria: ultimate load (Fu),
relative displacement maximum load (δmax) and connection stiffness (K). Figure 13 presents the
7
load – relative displacement curves of all push out tests and Table 3 lists the results obtained
for each connection system.
Figure 13 - Load-relative displacement behavior in the push out tests.
Table 3 - Push out tests results.
Bolted Bonded Bonded - bolted
Fu [kN] 88.0±6.0 127.5±5.3 158.6 ± 18.2
δmax [mm] 11.9±0.8 0.88±0.2 1.22 ± 0.2
K [kN/mm] 16.7±11.0 203.4±32.1 197.4 ± 32.7
The bonded-bolted shear connection system presented the highest ultimate load (158.6 kN).
Nevertheless, the stiffness of this connection system was slightly less than that presented by
the bonded connections: 197.4 kN/mm and 203.4 kN/mm respectively. However, if the standard
deviation values are taken into account the stiffness of those shear connection systems can be
considered to be virtually identical. This similarity is related to the much higher stiffness of the
bonded connection, when compared to the bolted connection. In terms of relative displacements
at maximum load, the bonded-bolted connections present an average value 38.9% higher than
that of the bonded connections. Nevertheless, in the former system, the displacements were
only 1.22 mm. With this regard, the bolted connections were the ones that presented the
highest values (11.9 mm), caused by the connections low stiffness (16.7 kN/mm). Even though,
this connection system failed for an average load of 88 kN, with a collapse mode that was the
most ductile amongst the three tested systems, followed by the bonded-bolted connection.
4.3 Flexural test
In order to study the flexural behavior of a GFRP–concrete hybrid beam a three-point flexural
test was conducted. The connection between the concrete slab and the GFRP profile was
materialized through a bonded connection, using the same epoxy adhesive, and an anchoring
0
25
50
75
100
125
150
175
0 5 10 15 20
Loa
d [
kN
]
Relative displacement [mm]
P1
P2
P3
C1
C2
C3
CP1
CP2
CP3
8
system that consisted of two sets of M10 two bolts located at a distance of 10.5 cm and 21.0 cm
from the beam´s extremity, in order to stop the interface cracking progression. The bolts were
designed to resist the normal stresses which developed along the concrete-adhesive interface.
To do so, it was considered that the normal stresses obtained by the FEM model were constant
along the section (conservative hypothesis).
The beam had a total length of 4.8 m and a span of 4.0 m. This test aimed to determine the
beam´s ultimate strength, flexural stiffness, failure mode and the influence and efficacy of the
proposed anchoring system. The geometry of the tested beam is presented in Figure 14 and in
Figure 15.
Figure 14 – Cross-section of the
hybrid beam (in mm)
Figure 15 – Geometry of the anchoring
zone (in mm).
4.3.1 Test set up
Figure 16 illustrates the test setup. The beam´s simply supported conditions were guaranteed
by placing a cylindrical hinge under each support, with one of the supports allowing for
longitudinal slinding. To avoid the lateral rotation of the beam, two metal plates were positioned
at a short distance from the profile´s web in the support sections. Load was applied
monotonically at the beam´s midspan, using a 300 kN hydraulic jack. Load was measured with
a Novatech load cell with a load capacity of 400 kN and longitudinal strains were measured in
section S using six strain gauges placed throughout the depth of the GFRP profile (Figure 17
and Figure 18). Bolt curvatures were also measured using strain gauges placed in diametrically
opposed sides and aligned with the beam’s longitudinal axis (Figure 19 and Figure 20).
9
Figure 16 - Flexural test setup.
Figure 17 – Flexural test setup frontal view. Figure 18 – Cross-section S.
4.3.2 Flexural test results
Figure 28 presents the load-mispan displacement behaviour of the GFRP-concrete hybrid
beam.
Figure 19 - Bolts (C1 to C8) and strain gauges (ε8 to ε23)
numbering.
Figure 20 - Longitudinal strain gauges in
the bolts.
10
Figure 21 - Load-midspan displacements behavior.
The GFRP-concrete hybrid beam presented a linear load-displacement behavior practically until
failure, which occurred for a load of 130.3 kN. As illustrated in Figure 28, the load-midspan
displacement behavior is in good agreement with the following equation, proposed in [13],
� = �� + �� =� ×
48 × ���� × �������
+ �� ×� ×
4 × �� × ��
where Ks is the proportion of shear carried by the GFRP profile, δf is the midspan displacement
due to the flexural component, δv is the midspan displacement due to shear, P is the load, L is
the beam span, EGFRP is the profile Young´s modulus, IeqGFRP
is the principal moment of inertia
of the homogenized section, GP the profile distortion modulus and Aw the profile web area.
In the tested beam, for service loads, the GFRP profile carried out 70% of the shear, with this
proportion increasing up to 85% prior to failure.
Failure occurred due to concrete debonding, which began in the beam´s right extremity and
rapidly developed until it reached midspan. Initially, debonding started a few millimeters inside
the concrete cover; however, at a distance of 4 to 8 cm beyond the second set of bolts,
separation occurred in the adhesive layer (Figure 22).
Figure 22 - Interface cracking after failure. Figure 23 - Bolts deformation after failure.
0
20
40
60
80
100
120
140
0 20 40 60 80 100 120
Loa
d [
kN
]
Midspan displacement [mm]
Composite
beam
GFRP profile
11
Among the six strain gauges used to monitor the axial deformations at midspan, only three (ε4,
ε5 and ε6) provided valid readings. Therefore, in order to estimate the neutral axis position, a
linear interpolation was required. This procedure showed that during the flexural test, in general,
the neutral axis was about 40 mm inside the concrete, which is in good agreement with the
cracked beam model proposed by Correia et al. [23].
Figure 24 – Axial strain vs. beam depth , for
various bending moments.
Figure 25 - Neutral axis vs load.
Prior to the test of the hybrid beam, a similar three-point flexural test had been conducted in the
GFRP profile [18]. Error! Reference source not found. compares the load-midspan deflection
behaviors of the hybrid beam and the GFRP beam, which failed for a total load of 60.2 kN.
Comparing the hybrid beam with the GFRP profile, it is possible to observe that the hybrid beam
still presents a linear behavior in service conditions. It is also possible to conclude that the
beam´s flexural stiffness increases 132.6 % and that the ultimate load increased 117.7 %.
These results show the efficiency of the concrete layer in terms of overall stiffness and strength
increase.
However, the proposed anchoring system proved to be inadequate in preventing the concrete
debonding. In fact, strain gauges showed that the applied bolts had a flexural behavior, instead
of an axial response.
5 Numerical analysis
5.1 Numerical analysis of the flexural test
The numerical analysis of the tested beam was divided in two parts. The first part consisted in
analyzing the serviceability behavior, similarly to what was done for Correia´s [13] beams (c.f.
section 3). Subsequently, a failure analysis was conducted, focusing in the interface cracking
evolution and in the stresses redistribution.
0
50
100
150
200
250
300
-2000 0 2000 4000 6000
Be
am
´s h
eig
ht
[mm
]
Axial strain[x10-6]
M = 10
M = 20
M = 30
M = 40
M = 50
M = 60
M = 70
M = 80
M = 90
200
220
240
260
280
300
0 20 40 60 80 100 120 140
Ne
utr
al
ax
is [
mm
]Load [kN]
NA experimental
NA cracked
NA uncrackedConcre
te
GF
RP
12
5.1.1 Serviceability analysis
Analyzing the stress state in the concrete slab at a depth of 0.5 mm for a load of 130.7 kN
(Figure 26), it is noticeable that for a serviceability state (for which an elastic behavior is
expected), bolts have practically no influence in the shear stresses distribution. The same does
not happen in the normal stresses distribution, where stress peaks are observed in the bolts
zones. These peaks have a maximum value of 1.3 MPa. In Figure 27 it is also observable that
the anchoring effects are merely local, hence having a very limited influence in the overall stress
distributions.
Figure 26 - Shear stresses in the tested beam for a
load of 130.7 kN.
Figure 27 - Normal stresses in the tested beam for
a load of 130.7 kN.
When observing the extremity of the beam it is possible to conclude that both normal and shear
stresses distributions are very similar to those of beam HB2 tested by Correia [13], which
means that there is no significant influence of the anchoring system.
The normal stress at the end of the beam is 15.34 MPa, which is 442% higher than fctm
(2.83 MPa). However, applying the same analysis as in section 3, namely extrapolating the end
values, the maximum stress is 5.65 MPa, which is approximately twice of fctm. If the average
stress criterion is applied, the maximum stress is only 35.9% higher than fctm. These results
demonstrate, once again, that the normal stresses at the end of the beam (obtained by a linear
elastic analysis) corresponding to the three criteria are higher than the average concrete tensile
resistance, which is in agreement with the observed cracking.
5.1.2 Failure analysis
Cracking propagation was simulated deleting adhesive elements from the beam´s extremity to
the studied sections. Due to the high stress concentrations near the beam axis (see Figure 27),
the interface crack was considered to propagate firstly between points Z=175 mm and
Z=225 mm, propagating to the entire interface after 5 mm, three parameters were analysed: (i)
longitudinal stress variations; (ii) transverse stress distribution; and (iii) anchoring stresses.
Longitudinally, this analysis has shown that both normal and shear stresses distributions suffer
strong variations as the crack propagates into the beam. In a first phase stresses increase and
then start to reduce at a distance of 10 mm from the beam extremity in the case of normal
stresses and at a distance of 50 mm in the case of shear stresses. When the crack is extended
Bolts
[MPa]
x [mm]
z [mm]
Bolts
[MPa]
z [mm]
x [mm]
Beam extremity
Beam extremity
13
to the second set of bolts, normal stresses become negative, which, according to continuous
mechanics, should imply that the crack would stop. This result suggests that this approach is
not adequate to study the debonding issue.
Along the transverse section of the beam, shear stresses concentrate mainly between Z =
175 mm and Z = 225 mm. However, as the cracking approaches the anchoring sections, shear
stresses start to concentrate near the bolts. In fact, the shear stresses in the bolts become
significant only after the complete cracking of the section occurs, after which they present an
approximately constant shear stress of -3.61 MPa in the first set bolts and 1.56 MPa in the
second set of bolts. The normal stresses distribution is very similar to that of the shear stresses,
concentrating mainly in the center of the profile, except in the anchoring sections, where
maximum stresses occur in the bolts.
Analyzing the shear and normal stresses in the bolts as the cracking propagates, it is obvious
that the highest stress values registered never attain the ultimate strength of the bolts. In fact,
the highest registered normal stress is -87.8 MPa, which is considerably less than 800 MPa, the
ultimate tensile stress of a M10 bolt (8.8 class). The same occurs in terms of shear stresses,
where maximum stresses are 13.7 MPa. This analysis demonstrates that the adopted anchoring
system is not efficient in stopping the profile debonding (as bolts mobilize only a small fraction
of their strength) nor in preventing the cracking propagation.
5.2 Parametric study
In order to determine how the interface stresses are influenced by different parameters, a
parametric study was conducted using the developed FEM. This analysis was made for beam
HB2, which had already been studied in section 3. The following six parameters were studied:
(i) concrete elasticity modulus; (ii) adhesive elasticity modulus; (iii) concrete slab thickness; (iv)
adhesive layer thickness; (v) free length at the end of the beam; and (vi) beam span.
The FEM analyses showed that, from the studied parameters, only span length does not
influence the interfacial stress distributions at the extremity of the beam.
The beam free length proved to have a major influence in the interfacial stresses at the
extremity of the beam. In fact, for a 800 mm long free length, the interfacial stresses are
approximately 5.7 times smaller compared to a free length of 200 mm. Also if there is no free
length, normal stresses are virtually inexistent. The concrete layer thickness also proved to
have a major influence in these stresses. In this case, a 100 mm thick concrete layer is
associated to stresses that are 3 times higher than those corresponding to a 50 mm thick layer,
and 1.3 times higher than those corresponding to a 150 mm thick layer. Figures 28 and 29
show both normal and shear stresses variations for the concrete layer thickness (where Hc
corresponds to the reference thickness of 100 mm) and for different free lengths (where D
corresponds to the reference free length of 400 mm).
14
Figure 28 - Stress variations for different concrete
layer thickness´s.
Figure 29 - Stress variations for different free
lengths.
An increase in the concrete elasticity modulus leads to an increase of the interfacial stresses.
Although this variation is very small for normal grade concretes, it can have a significant effect if
high-strength concretes are used, as they become more prone to interface cracking. The
variation of the adhesive elasticity modulus influences the interfacial stresses in a logarithmic
way. Hence, using a polyurethane adhesive (with a low elasticity modulus) it is possible to
reduce the interface stresses by 50%, compared to the epoxy adhesive used in the
experiments. Finally higher adhesive layer thicknesses cause the extremity stresses to diminish,
as they create a more flexible connection between concrete and the GFRP profile.
6 Conclusions
In this work the flexural behavior of GFRP-concrete hybrid beams with a bonded shear
connection system was investigated. The finite element models developed in the numerical
study allowed to obtain a detailed description of the interfacial stresses at the beam extremity,
where debonding failure mechanism is triggered. These models have demonstrated that prior to
failure, normal stress peaks, several times higher than concrete average tensile strength, occur
near the beam extremity and around the beam´s longitudinal axis, thus causing concrete
cracking. In order to be able predict the failure load, three normal stress criteria were assessed:
(i) maximum local stress; (ii) maximum extrapolated stress; and (iii) average stress. All these
criteria, which compared normal stresses with the value of fctm, have proved to be too
conservative, with the best predictions being provided by the average stress criterion.
An anchoring system was developed to prevent concrete cracking from spreading. This system,
consisting of two sets of M10 bolts placed at a distance of 10.5 cm and 21.0 cm from the
extremity of the beam, was primarily tested in push out tests and compared with simply bolted
and simply bonded connections. In spite of having a similar stiffness, this system has shown a
much more ductile failure than the simply bonded connections and the highest ultimate strength
amongst the three systems. However, during the flexural test, this system has shown to be
0,0
0,2
0,4
0,6
0,8
1,0
1,2
0,5 1,0 1,5 2,0 2,5
σi
/ σ
Hci / Hc
Shear
Normal
-0,5
0,0
0,5
1,0
1,5
0,0 0,5 1,0 1,5 2,0
σi/
σ
Dai / Da
Shear
Normal
15
unable to stop the interface cracking. The flexural test demonstrated that the composite beam
has a linear behavior in serviceability state, suffering a small stiffness reduction only
immediately before failure.
The numerical model of the tested beam proved that the anchoring system with steel bolts had
only a local effect and that their maximum stresses are only a small fraction of the material
ultimate strength. The FEM was also used to determine the normal and shear stresses
distributions as the cracking developed into the beam. This analysis showed that when the
crack propagates, the normal stresses at the second set of bolts become negative.
Nevertheless, test showed that cracking continues to develop, hence proving that the
continuous mechanic theory may not be suitable to explain the debonding mechanism. With this
regard, a fracture mechanics approach may provide better results.
The parametric study investigated the influence of the elasticity modulus and thickness of both
concrete and adhesive, the free length and the beam span. From the analyses performed it was
possible to conclude that longer beams are less susceptible to the debonding failure as the
interfacial stresses remain unchanged with the variation of the span. It was also possible to
verify that the careful choice of materials and geometries can considerably reduce interfacial
stresses.
16
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