TENSION OF REINFORCEMENT BARS EMBEDDED IN CONCRETE PRISMS STRENGTHENED WITH CFRP PLATES

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Nov 26, 2013 (3 years and 8 months ago)

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FRPRCS-9 Sydney, Australia Monday 13 – Wednesday 15 July 2009
1

TENSION OF REINFORCEMENT BARS EMBEDDED IN
CONCRETE PRISMS STRENGTHENED WITH CFRP PLATES

Piotr RUSINOWSKI
1,2
Fedja ARIFOVIC
1,3
Björn TÄLJSTEN
1


1
Department of Civil Engineering, Technical University of Denmark, Lyngby, Denmark
2
Norut Northern Research Institute AS, Narvik, Norway
3
Alectia A/S, Virum, Denmark

Keywords: tension stiffening, crack spacing, FRP, strengthening.

1 INTRODUCTION

Fibre Reinforced Polymers, FRP, offer excellent corrosion resistance to environmental agents as
well as the advantages of high stiffness-to-weight and strength-to-weight ratios when compared to
conventional construction materials. For this reason FRP composites are becoming a material of
choice in an increasing number of rehabilitation and retrofitting projects around the world. The
application of bonded strengthening requires non-conventional design issues. One such design issue
is the debonding problems in externally bonded FRP strengthening applications that have been a
concern and a research challenge since the initial development stages of the strengthening method.
Debonding failures of tension face plated beams are crack induced and thus the cracking
behaviour of strengthened beams is of vital interest. For this reason it is important to study the crack
formation and crack distribution in strengthened concrete members. This paper presents the
experimental and analytical study of externally strengthened concrete prisms with an embedded steel
rebar subjected to tension. Stress distribution in such structures has been studied elsewhere, both
experimentally [1] and numerically [2]. The present research is focused on optical measurements of
cracking and assessment of a simple analytical model predicting crack spacing.

2 EXPERIMENTAL STUDY

A total number of 15 tension specimens were tested in the experimental programme. The
specimens consisted of single steel reinforcement bar embedded in a concrete prism. The specimens
were cast from three batches consisting of six prisms, in each case the same recipe for concrete was
used. In each batch three different types of internal reinforcement were used, as shown in Figure 1a.
One pair of prisms was reinforced with two short rebars remaining a 200-mm-long gap in the centre.
Other prisms were reinforced with continuous steel rebars weakened by grinding in the middle section,
this in order to provoke yielding the rebar. Apart from the longitudinal rebar all specimens were
internally reinforced with steel spirals in order to delay the splitting failure.


Fig. 1 General design of tested beams and location of strain gauges
FRPRCS-9 Sydney, Australia Monday 13 – Wednesday 15 July 2009
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Table 1 Test matrix.
Specimen #

φ20 cont. φ16 cont. φ16 gap
f
cc

[MPa]
f
ct

[MPa]
E
c

[GPa]
t
f
[mm]
E
f

[GPa]
CFRP type
Batch 1 1-1A, 1-1B 1-2A, 1-2B 1-3A, 1-3B 52.8 2.8 36.7 1.4 155.0 StoFRP Plate E 100 C
Batch 2 2-1A, 2-1B 2-2A, 2-2B 2-3A, 2-3B 51.8 2.9 36.3 1.4 260.0 StoFRP Plate M 100 C
Batch 3 3-1 3-2 3-2 49.6 2.7 35.6 - - -
f
cc
– concrete compressive strength, f
ct
– concrete tensile strength, E
c
– concrete modulus of elasticity, t
f
– plate thickness,
E
f
– laminate modulus of elasticity

After curing of concrete, two opposite sides of each prism from batches 1 and 2 were sandblasted,
properly cleaned and primed. The specimens were then strengthened with CFRP plates with high
tensile strength (batch 1) and high modulus of elasticity (batch 2). Specimens from batch 3 remained
without strengthening and three of them were chosen to testing. Half of the strengthened specimens
were clamped at the end of the concrete section in order to examine influence of splitting on crack
pattern and crack widths.
Actual concrete and steel properties were determined in small scale tests. Concrete compressive
strength of each batch was determined by testing cylinders (100×200 mm) and the tensile strength by
testing wedge splitting samples (100×100×100 mm). The actual properties of CFRP plates were not
determined and material properties provided by the manufacturer were used in further analysis. Test
matrix including material properties is presented in Table 1.
All test specimens were loaded by applying a tension force to the steel rebar. Measurements on
tested specimens during loading were mainly focused on monitoring of development of crack pattern.
This was performed with use of optical equipment Aramis provided by GOM mbH. The equipment
consists of two digital cameras that take photos in specified intervals. The measured area must be
covered with a spatter pattern so that each point can be recognized by the software. Each picture is
compared with the initial state and detected movements are recalculated into strains. The system,
however, is not accurate enough to measure strains in brittle materials like concrete but is very
convenient in measuring crack distribution and crack widths. Apart from the optical monitoring
elongation of prisms and the strains in CFRP plates and steel rebars were measured. The location of
LVDTs and strain gauges is shown in Figure 1.

3 ANALYTICAL STUDY

Analyses of concrete prisms with the internal rebars subjected to tension may be found in many
reports and compilation of several papers may be found in [3]. Stabilised crack spacing for tested
specimens is estimated with the formula recommended in CEB-FIP model code [4], Eq. 1.

φ
φ
φ
τ
ρ ρ ρ
= ⋅ = ⋅ =
2 2
0.185
3 2 3 3.6
ct
m
b
f
s
(1)

where, s
m
– stabilised crack spacing,
φ
– diameter of steel rebar,
τ
b
– bond stress between concrete
and steel,
ρ
s
– reinforcement ratio (A
s
/A
c
).
Analyses of externally strengthened prisms have not been found in any building codes or design
guidelines. A numerical model is presented in [5]. The analysis begins with minimum crack spacing
estimation by guessing width of the first crack and load when the second crack arises. Strain or stress
analysis can be then performed and possible debonding can be detected.
In the present study a simple equation for the average crack spacing is given. It is assumed that
first cracks appear when the stresses in concrete exceed the effective tensile strength of concrete.
The effective tensile strength is the tensile strength of cracked concrete. Usually it may be taken as
the half of the tensile strength of the uncracked concrete. It is further assumed that the FRP-plate
governs the crack formation in the same way that internal steel bars do in non-strengthened
structures. Thus, when the forces transferred between a tensile bar and the concrete and between the
FRP-plate and the concrete are such that they exceed the effective tensile strength of the concrete,
the average crack spacing may be written as:

λ
πφ λ
= ⋅
+
4
3
c
m
s s f f f
A
s
n n b
(2)

In this equation the quantities are A
c
– cross-sectional area of the prism b
f
– plate width,
λ
s
– ratio
between the maximum transferable stress in steel/concrete interface and concrete tensile strength,
FRPRCS-9 Sydney, Australia
Monday 13 – Wednesday 15 July 2009
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λ
f
– ratio between the maximum transferable stress in FRP/concrete interface and concrete tensile
strength, n
s
– number of rebars, n
f
– number of FRP-plates.
The approach here is to define maximum stress the FRP-plate and the steel bar may transfer to
the concrete. Thus, the rate of the maximum transferable stresses at the interfaces (concrete to rebar
and FRP-plate to concrete) in relation to the effective tensile strength of concrete is introduced through
λ
-factors (
λ
s
and
λ
f

factors for the steel bar and the FRP-plate, respectively). In non-strengthend
structures, on basis of tests, the maximum transferable stress at the steel bar to concrete interface
was found to be approximately half the effective tensile concrete strength (i.e.
λ
s
= 1.8, [4] or
λ
s
= 2,
[6]). It is likely that this relation changes when the crack spacing is influenced by the FRP-
strengthening. In this study new
λ
-values are suggested for such cases.

4 RESULTS

The experimental study was focused mainly on optical measurements of the crack distribution and
influence of clamping applied to avoid splitting failure. Among the unclamped strengthened specimens
only one did not fail by splitting. Spiral reinforcement at the ends of the prisms was sufficient to
prevent from splitting failure in specimen 2-1A with continuous
φ
16 rebar and CFRP of lower
E-modulus. In this case as well as for all clamped specimens steel yielding outside the prisms
occurred. It is worth noting that splitting crack initiated also in all clamped specimens.
Figure 2 shows strain-load distributions in all strengthened specimens where strains were
measured in the CFRP plates in the mid-section. Clamped specimens reached generally higher loads
as they were prevented from splitting and thus their load capacity was limited only by the yield
strength of the rebars. It was supposed that applying too much tension to the bolts in the clamps may
restrain slippage of CFRP plates at the ends of the prisms and influence strains in cracking stage.
Apart from the ultimate loads, however, no significant difference between the test results for clamped
and unclamped specimens can be noticed.
Optical equipment Aramis was a useful tool in measurement of crack widths and crack spacing.
Considering high strengthening ratio, cracks were very difficult to notice with bare eye due to their very
small widths. As an example, Figure 3 shows development of crack pattern in specimen 2-2A as a
sequence of pictures obtained from Aramis at different load levels. With help of this data, estimation of
crack spacing and load when stabilised cracking occurs may be done easily. It appears that cracks are
relatively straight and distributed evenly. In some specimens, however, inclined cracks initiating
debonding were observed.
Table 2 shows average crack spacing measured in the experiments and estimated with Eq. 2. It
appears that the analysis estimates significantly smaller crack spacing. In the strengthened specimens
factor
λ
f
=
λ
s
= 1 is applied to achieve the best correlation between the tests and results of Eq. 2. In

Table 2
Crack spacing measurements vs. theory.
1-1 1-2 1-3 2-1 2-2 2-3 3-1 3-2
λ
s
/ λ
f

1.0 / 1.0 1.0 / 1.0 - / 1.0 1.0 / 1.0 1.0 / 1.0 - / 1.0 1.8 / - 1.8 / -
s
m,theory
[mm]
51 53 67 51 53 67 118 147
Unclamped (A)
62 60 102 69 63 99 155 221
s
m,exp
[mm]
Clamped (B)
70 59 66 60 71 81 - -
Unclamped (A)
82 89 65 74 85 67 76 67
exp.
,
,
100%
m theory
m
s
s

Clamped (B)
72 90 101 85 75 81 - -

0 0.05 0.1 0.15 0.2 0.25
Strain [%]
0
50
100
150
200
Load, P [kN]
φ20 cont.
E
f
=155 GPa
clamped
unclamped
φ16 cont.
E
f
=155 GPa
unclamped
clamped
0 0.05 0.1 0.15 0.2 0.25
Strain [%]
0
50
100
150
200
Load, P [kN]
φ20 cont.
E
f
=260 GPa
clamped
unclamped
φ16 cont.
E
f
=260 GPa
unclamped
clamped
0 0.05 0.1 0.15 0.2 0.2
5
Strain [%]
0
50
100
150
200
Load, P [kN]
φ16 (gap)
E
f
=260 GPa
clamped
unclamped
φ16 (gap)
E
f
=155 GPa
unclamped
clamped

a)
b)
c)

Fig. 2
CFRP strains in the mid-section, a) specimens 1-1 and 1-2, b) specimens 2-1 and 2-2, c)
specimens 1-3 and 2-3
FRPRCS-9 Sydney, Australia
Monday 13 – Wednesday 15 July 2009
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Fig. 3
Crack development in specimen 2-2A (
φ
16, E
f
= 260 GPa)

non-strengthened specimens
λ
s
= 1.8 is used. Furthermore, it is shown in Table 2 that there is no
significant difference between crack spacing for clamped and unclamped specimens and such type of
clamping may be used in future experiments.

5 CONCLUSIONS

Simple model for calculating average crack spacing for CFRP strengthened concrete specimens
loaded in tension is introduced on basis of an empirical determination of the maximum transferable
shear stress at the FRP-plate to concrete and steel bar to concrete interfaces. The magnitude of such
stresses is defined through a
λ
-factor that is the rate of the maximum transferable stress in relation to
the effective tensile strength. In non-strengthened specimens this factor is approximately set to 2. By
the use of
λ
= 1.8 the average crack spacings were calculated slightly lower than measured on non-
strengthened specimen reported here. The correlation was thus reasonably good.
However, for the strengthened specimens, both
λ
s
and
λ
f
were set to 1 in order to achieve
reasonable correlation to measured crack spacings. It was thus shown that the application of the FRP-
plate to tensile specimens increases the maximum transferable stresses at the interfaces. The reason
for this may be the fact that in the small specimens with a high strengthening degree the influence of
the FRP-plate is greater than in RC beams, for instance. Furthermore, in small specimens the
effective tensile strength exhibits plastic behaviour. It is therefore also likely that the application of the
non-changed
λ
-values (i.e. equal to 1.8 or 2.0) may be appropriate in the case of tension face FRP-
plated beams.

ACKNOWLEDGEMENTS

The research programme was financed by The Norwegian Research Council through the strategic
institute programme RECON at Norut Northern Research Institute AS.

REFERENCES

[1] Ueda, T., Sato, Y., Yamaguchi, R. and Shoji, K., “Study on behavior in tension of reinforced
concrete members strengthened by carbon fiber sheet”, Journal of Composites for
Construction 6(3), 2002, pp 168-174.
[2] Ferretti, D. and Savoia, M., “Non-linear model for R/C tensile members strengthened by FRP-
plates”, Engineering Fracture Mechanics 70(7-8), 2003, pp 1069-1083.
[3] Elfgren, L. and Noghabai, K., eds., Tension of Reinforced Concrete Prisms - Round Robin
Analysis and Tests on Bond - Report of RILEM TC 147-FMB, Research Report, Luleå
University of Technology, 2001
[4] CEB-FIP, Model Code 1990. Design Code, Comité Euro-International du Béton and
Féderation International de la Precontrainte, Thomas Telford, London, 1993, 437 pp.
[5] Liu, I. (2005), Intermediate Crack Debonding of Plated Reinforced Concrete Beams, PhD
thesis, The University of Adelaide.
[6] Nielsen, M. P., Beton 1 Del 3, 2
nd
ed., Department of Civil Engineering, Technical University of
Denmark, 2005, (in Danish).