Open Journal of Civil Engineering, 2012, 2, 115-124 Published Online September 2012 (
Flexure-Compression Testing of Bridge Timber Piles
Retrofitted with Fiber Reinforced Polymers
Pablo Caiza, Moochul Shin, Bassem Andrawes
Department of Civil and Environmental Engineering, University of Illinois at Urbana-Champaign, Urbana, USA
Received July 18, 2012; revised August 20, 2012; accepted September 4, 2012
The adequacy of using Fiber Reinforced Polymer (FRP) retrofit technique to restore the flexure-compression behavior
of deteriorated bridge timber piles is examined experimentally in this paper. Sixteen specimens are tested monotonically
under eccentric compressive loading. The specimens are first tested in their unretrofitted condition to determine their
elastic properties. Each specimen is then cut and connected (posted) using the proposed FRP retrofit technique, and re-
tested. The results show that the retrofitted specimens are capable of reaching same or higher strengths than that of the
unretrofitted specimens with minimal reduction in their stiffness. Based on the experimental results, a design equation is
presented to compute the volumetric ratio of FRP needed for retrofitting bridge timber piles under eccentric load.
Keywords: Bridges; Timber Piles; Eccentric Loading; Fiber-Reinforced Polymer (FRP)
1. Introduction
The relatively limited maintenance and attention which
small rural bridges often receive throughout their service
life could potentially result in the rapid deterioration of
these bridges’ structural conditions. Due to this fact along
with the consistent increase in population and loading
demands, many of these bridges reach a point where they
fail to accommodate current traffic volumes, vehicle
sizes, and weights [1-3]. Taking into account the high
cost of replacing those bridges, every effort is needed to
develop sound and cost-effective retrofit measures that
can effectively extend the service life of these bridges.
Among the bridges that are commonly found in rural
areas are the ones that are supported on timber piles.
Bridges supported on timber piles are vulnerable to ex-
treme loadings as well as harsh environmental conditions
which could lead to rapid decay of the wood if the piles
are not regularly monitored [4]. In the United States and
other parts of the world, some measures are often taken
to delay or prevent the decay of bridge timber piles, such
as the use of creosote oil and the regular inspection of the
piles. If inspection reveals that a portion of the pile is in a
bad condition and in need for replacement, a technique
often known as “posting” is utilized. Posting a pile is a
process where a portion of the pile is cut out and replaced
with a new piece of timber, see Figure 1. This new piece
is connected to the existing pile by embedded metal fas-
teners [5]. The post connection at the interface between
the existing timber and the new timber is capable of
transmitting axial compression forces only; however it is
not designed to transfer moment. Nonetheless, due to its
cost-effectiveness this type of retrofit technique is pre-
ferred to a complete replacement of the entire pile or
encasing of the timber pile in a thick reinforced concrete
Although the pile posting retrofit technique is com-
monly used, its limited ability to restore the flexural ca-
pacity of the piles is considered to be a major drawback
that could potentially cause severe consequences as in
the case of a relatively recent bridge collapse incident in
the State of Illinois [6,7]. Investigation suggested that the
collapse of this bridge was due to the presence of flexural
bending in addition to axial compression applied to the
posted timber piles. Hence, there is a dire need for a sim-
ple and reliable procedure that could enhance the flexure-
compression behavior of posted piles at the post connec-
tion interfaces. One of the relatively recent approaches
studied to restore the load-carrying capacity of heavily
Figure 1. Replacement of a wood piece from a damaged
timber pile.
opyright © 2012 SciRes. OJCE
decayed timber piles is the use of fiber reinforced poly-
mer (FRP) wraps on the top of a relatively thin cementi-
tious grout shell around the timber pile. In most of the
research done in this area, glass-FRP (GFRP) wraps are
basically used to confine the grout shell and the timber
core, in such a way that only the axial compression ca-
pacity of the piles is improved [8]. However, no studies,
to the knowledge of the authors, have been conducted
specifically to investigate the combined flexure-com-
pression behavior of timber pile post connections retro-
fitted with FRPs. This paper aims at studying the feasi-
bility of using FRP wraps as a mean for splicing the new
post with the existing pile in a way that ensures the flex-
ural integrity of the retrofitted pile. The paper discusses
an experimental program which was carried out on retro-
fitted timber piles tested under uniaxial concentric and
eccentric loading. The paper starts by discussing the
proposed retrofit technique in Section 2 followed by a
description presented in Sections 3 and 4 for the timber
pile specimens used in the tests and the test setup
adopted in this study, respectively. A detailed discussion
is then presented in Section 5 on the approach used in
designing the FRP wraps used in the tests followed by a
discussion in Section 6 on the test results. Finally, the
test results are used in Section 7 to develop design rec-
ommendations for FRP-retrofitted timber piles subjected
to eccentric loading.
2. Proposed Retrofit Technique
Figure 2 shows schematics of the proposed retrofit tech-
nique which was adopted in this study. The first step is to
align the timber post and the existing pile and connect
them using a vertical steel drift pin at the center of the
round pile (Figure 3(a)) or by using nails driven at the
sides of the pile (Figure 3(b)). After the two spliced
pieces of timber are attached, the round surface of the
timber pile is sanded, and a thin mortar shell (less than 2
mm) could be applied to provide an evenly round surface
for the FRP sheets which will be wrapped in the next step.
Another advantage of using the mortar shell is to block
the creosote oil that could possibly be within the wood
from contaminating the FRP resin. Finally, the FRP
sheets are applied around the pile with resin using the
hand lay-up process.
3. Specimen Description
The experimental specimens used in this study were cut
from pile samples that were retrieved from several
bridges in the state of Illinois. The retrieved samples
were different in their dimensions (length and diameter)
and their age, which directly affected the level of dete-
rioration in each specimen. It should be noted, however,
that all timber piles sustained relatively moderate levels
of decay. Figure 4 shows different types of decay which
was observed in the tested specimens including vertical
cracks (Figure 4(a)), partially crushed fibers (Figure
4(b)), irregular shape of the section due to the existence
of a knot (Figure 4(c)), and cracks filled with mold and
fungus (Figure 4(d)).
Figure 2. Proposed retrofit technique.
(a) (b)
Figure 3. Detail of a post connection of a timber-pile.
(a) (b) (c) (d)
Figure 4. Different types of decay observed in the tested pile specimens: (a) Vertical crack; (b) Partially crushed fibers; (c)
Irregular shape; (d) Mold and fungus inside of cracks.
Copyright © 2012 SciRes. OJCE
In total, 16 specimens were prepared by cutting them
from the retrieved piles. The specimens were cut to a
constant length of 1220 mm, and their diameter varied
between 249 mm and 318 mm. Each specimen was tested
twice under combined flexure-compression loading. The
purpose of the first test was to determine the strength and
stiffness of the unretrofitted timber piles. Next, the
specimen was cut perpendicular to its longitudinal axis in
two equal halves, and the two halves were spliced to-
gether using the proposed retrofit technique, i.e. with a
steel drift pin or nails and FRP sheets. The second flex-
ure-compression test was carried out next, to examine the
behavior of the spliced specimens after being retrofitted
with FRP wraps. The specimen test matrix is summarized
in Table 1 . The nomenclatures used herein for the unret-
rofitted and retrofitted specimens are SPxU and SPxR,
respectively, where “x” indicated the specimen number.
Two of the 16 specimens were retrofitted with carbon
FRP (CFRP) sheets (SP9 and SP13), while the rest of the
specimens were wrapped with GFRP sheets.
In order to examine the effect of using mortar shell
prior to wrapping the FRP sheets, a mortar shell was
utilized in the retrofit of specimens SP8, SP9, SP10, and
SP11. To explore the effect of the presence of excessive
decay and cavities in the piles, a wedge was introduced
in two of the tested specimens (SP12 and SP13) as
shown in Figure 5. As shown in the figure, the wedge
was introduced by cutting out a portion of the pile and
filling it with mortar prior to FRP wrapping. Further-
more, since the use of center-lined drift pins to align the
two spliced timber pieces could be problematic in the
field, the possibility of using diagonally driven nails at
Table 1. Specimens test matrix.
Specimen D [mm] Retrofit Details
SP1 262 GFRP
SP2 290 GFRP
SP3 249 GFRP
SP4 282 GFRP
SP5 262 GFRP
SP6 277 GFRP
SP7 254 GFRP
SP8 318 GFRP, mortar shell
SP9 305 CFRP, mortar shell
SP10 315 GFRP, mortar shell
SP11 310 GFRP, mortar shell
SP12 318 GFRP, wedge
SP13 292 CFRP, wedge
SP14 259 GFRP, nails
SP15 290 GFRP, nails
SP16 318 GFRP, nails
the sides of the specimen was examined in three speci-
mens (SP14, SP15, and SP16). Figure 6 presents pictures
illustrating the procedures used in preparing the tested
Table 2 summarizes properties and details of the ma-
terials used in retrofitting the specimens. Vinylester resin
was used to form the matrix of both GFRP and CFRP.
The glass fabric used was woven roving, and was se-
lected for its relatively inexpensive, high impact, and
high strength two-directional reinforcement properties.
The carbon fabric used was plain weave, which is char-
acterized with light-weight and high stiffness. The con-
crete mortar has a relatively high compressive strength of
22 MPa at 1 day, as well as a fast final set time of 2
hours. It comprises one component, microfiber, silica
plus polymer mortar.
Figure 5. Wedge at the post connection.
(a) Pile cutting (b) Steel pin installation
(c) Sanding (d) Mortar shell
(e) FRP wrapping
Figure 6. Details of the pile retrofit procedure.
Copyright © 2012 SciRes. OJCE
Table 2. Properties of the materials used to retrofit the
timber piles.
Compression force
Material Properties
Glass FRP Modulus of elasticity =12,400 MPa,
thickness = 0.6 mm
Carbon FRP Modulus of elasticity = 135,800 MPa,
thickness = 0.3 mm
Concrete mortar Compressive strength (1 day) = 22 MPa,
final set time of 2 hours.
4. Test Set-Up
The tests performed in this study were conducted using a
2670 kN MTS uniaxial servo-controlled compression
machine. The actuator was controlled by an INSTRON
8800 controller. The testing frame contained an internal
load cell and an internal Linear Variable Differential
Transducer (LVDT) to measure actuator position. The
specimens were instrumented with two vertical exten-
someters placed at the mid-height of the specimen on two
opposite sides to measure axial strains (see Figure 7).
Two additional LVDTs were placed at the top and
bottom of the pile specimen to monitor the relative dis-
placement between the specimen and the end plates dur-
ing testing. The typical disposition of the LVDT located
at the top of the pile is also shown in Figure 8.
Special fixture plates were designed, manufactured,
and mounted at both ends of the specimen to apply the
axial load eccentrically. The specimens were bolted to 38
mm thick steel plates on each end using ten 19 × 151 mm
bolts for the top plate and eight 19 × 151 mm bolts for
the bottom plate. Figure 9 shows schematics of the spe-
cial plates used for applying the eccentric load.
The plates were loaded through 51 mm diameter roll-
ers which were placed at distance of 76 mm from the
centroidal axis of the pile specimen (see Figure 10).
Figure 11 shows an installed specimen in the 2670 kN
machine frame. As indicated by the described test setup,
Figure 7. Longitudinal extensometers on both sides of the
Tension sideCompression side
Figure 8. Installed LVDT at the top of the pile.
Figure 9. Steel fixture plates used to apply eccentric axial
load on the specimens (Dimensions in mm).
Figure 10. Typical roller at the base of the pile.
Figure 11. Installed specimen in the MTS testing machine.
Copyright © 2012 SciRes. OJCE
a monotonic compression load with an eccentricity of 76
mm was applied to the timber-pile specimens. The tests
were conducted under displacement control at 1.27 mm
per minute cross-head rate. First, an unretrofitted speci-
men was tested until a point where nonlinear behavior of
the specimen was observed. The maximum load recorded
at this point was considered as the elastic strength of the
pile specimen. Then, the specimen was cut to simulate a
post connection, and retrofitted with a proper number of
FRP wraps satisfying the elastic design criteria of the
FRP retrofit, which is discussed in the following section.
The retrofitted specimen is then reloaded until the first
sign of nonlinearity in the force-deformation relationship
is observed.
5. Elastic Design of FRP Wraps
In order to determine the number of FRP layers needed
for the retrofitting of the piles, an elastic design proce-
dure was adopted. The procedure was based on the as-
sumption that at the interface section between the post
and the old pile, the tensile stresses are resisted entirely
by the FRP, and the compressive stresses are resisted by
both wood and FRP. At this section, the portion of the
wood resisting compression is called “effective area”
(see Figure 12).
To determine the number of FRP layers that gives the
same flexural stiffness in the retrofitted pile as that of the
unretrofitted pile, the following formula was used:
Wood Wood
EI (1)
where Ewood is the modulus of elasticity of wood, I is
moment of inertia of the complete circular section of the
original pile, Icomp is the moment of inertia of the effec-
tive area of the pile section (Figure 12), EFRP is the
modulus of elasticity of FRP, and
is the moment
of inertia of the FRP wraps used for the retrofitting of
timber pile.
Equation (1) was used to derive the formula used to
compute the required FRP volumetric ratio, ,
is defined as the ratio between the volumes of FRP and
timber. The volumetric ratio of FRP can be expressed as
2Dnt D
, where D is the diameter of the pile, n
(a) Original section (b) Retrofitted section
Figure 12. Original vs. retrofitted section at the post con-
nection interface.
is the number of FRP layers, t is the thickness of a single
FRP layer. Using Equation (1), the volumetric ratio of
FRP can be described as:
111comp Wood
 
Additionally, Equation (2) can be rewritten in terms of
the required number of FRP layers as follows:
11 1
comp wood
 
The moment of inertia of the effective area of the tim-
ber pile section is calculated from the following equa-
*2 *221
sin ;
yR R
 
where c is the depth of the effective area of the timber-
pile section (Figure 12). Note that c can be calculated
from the classical equation for elastic stresses [10,11]:
c (5)
where, A is area of the circular section of the pile, and e
is eccentricity.
Figure 13 shows an example of the strain distribution
at the interface/mid-height determined from the tests us-
ing the two extensometers attached on both sides of the
specimen. In this figure, compressive strains are consid-
ered negative and tensile strains positive. The presented
distributions are for the unretrofitted (SP4U) and retro-
fitted (SP4R) specimen 4. Using the experimental strain
data, the depth of the effective area was found to be 201
mm and 218 mm for SP4R and SP4U, respectively with a
difference of 7.8%. For comparison, the analytical depth
of the effective area was also computed using Equation
(a) SP4R (b) SP4U
Figure 13. Strain distribution at mid-height of specimen
SP4 (D = 282 mm).
Copyright © 2012 SciRes. OJCE
Copyright © 2012 SciRes. OJCE
(5) and was found to be 206 mm; hence the differences
between the analytical and experimental c values of the
retrofitted and unretrofitted specimens were approxi-
mately 2.5% and 5.8%, respectively. This minor differ-
ence could possibly be due to imperfections in the piles’
circular section and decay and damage of the wood.
since it is quite close to the mean value and within 95%
confidence interval of the experimental values.
Based on the described design procedure, the number
of FRP layers used for each specimen was calculated and
is summarized in Table 3. The second row in the table
corresponds to the number of FRP layers predicted using
Equation (3), while the third row corresponds to the
number of layers actually used in the tests. The fourth
row shows the FRP volumetric ratio (
As stated earlier, the diameter of the section of the
timber piles (D) varied between 249 and 318 mm, corres-
pondingly comp
). As shown in
the table, a number of layers similar to or greater than the
predicted number from Equation (3) was utilized for ret-
rofitting specimen SP1 and specimens SP8 through SP13.
However, since most of these tests showed that the retro-
fitted specimens reach higher strength than that of the
unretrofitted specimens, and in order to optimize the de-
sign of the FRP sheets, using smaller number of layers
than those computed based on Equation (3) was also con-
sidered. It is important to point out that the CFRP volu-
metric ratios used for SP9 and SP13 are much smaller
than those of the GFRP retrofitted specimens due to the
significantly higher modulus of elasticity of CFRP com-
pared to GFRP.
varied between 0.55 and 0.61. If the
mean value for comp
I (i.e. 0.58) is used, Equation (2)
can be further simplified to
1 1
FRP 0.42 Wood
  (6)
Equation (6) shows that the ratio wood
E is the key
factor governing the design of FRP, but the modulus of
elasticity of wood and FRP can vary substantially [12].
As shown in Tab le 2, the modulus of elasticity of GFRP
and CFRP used in this study was 12,400 MPa and
135,800 MPa, respectively. The modulus of elasticity of
the wood on the other hand was determined experimen-
tally using the results of the tests conducted on unretro-
fitted specimens. Based on the results of the 16 tested
specimens, the mean value of the modulus of elasticity of
wood was found to be 4474 MPa with standard deviation
of 1234 MPa. In addition, there is a 95% confidence that
the value of the modulus of elasticity is between 3820
MPa and 5136 MPa. It is worth noting that the National
Design Specification for Wood Construction [13] rec-
ommends the following values for modulus of elasticity
of treated round timber piles for normal load duration
and wet service conditions of Northern and Southern Red
Oak: E = 8618 MPa and Emin = 4550 MPa. To be on the
conservative side, the value of Emin was chosen as the
modulus of elasticity for timber in the pile retrofit tests
6. Experimental Results
The results of the flexure-compression tests of all 16 pile
specimens are summarized in Table 4. The table is di-
vided into four sections depending on the retrofit details
used for the specimens. The specimens were retrofitted: 1)
with FRP sheets without mortar shell, 2) with FRP sheets
and mortar shell, 3) with FRP sheets and mortar-filled
wedges, and 4) with only FRP sheets but the pile was
connected using nails while other specimens under the
previous sections were connected using steel drift pins
(see Figure 3). The second and third rows in each section
of the table represent the maximum compression forces
of the unretrofitted (unret ) and retrofitted (ret ) speci-
mens, respectively. The fourth row presents the ratio
between retrofitted and unretrofitted maximum forces. It
Table 3. Number of FRP layers and volumetric ratios.
No. of FRP layers Equation (3) 8 9 2 9 9 9 2 9
Test 8 9 2 10 9 9 2 7
ρFRP Test 0.079 0.074 0.008 0.084 0.076 0.073 0.008 0.062
No. of FRP layers Equation (3) 8 9 8 8 8 8 9 9
Test 6 7 6 6 6 6 6 7
ρFRP Test 0.062 0.064 0.059 0.056 0.061 0.060 0.053 0.057
Table 4. Maximum forces of unretrofitted and retrofitted specimens under flexure-compression tests.
Without mortar shell
Specimen SP1 SP2 SP3 SP4 SP5 SP6 SP7
Punret [kN] 347 436 289 863 516 565 325
Pret [kN] 387 529 329 912 467 565 3253
Pret/Punret 1.12 1.21 1.14 1.06 0.91 1.00 1.00
With mortar shell
Specimen SP8 SP9 SP10 SP11
Punret [kN] 1406 787 555 681
Pret [kN] 1397 778 681 810
Pret/Punret 0.99 0.99 1.23 1.19
With wedge
Specimen SP12 SP13
Punret [kN] 592 725
Pret [kN] 801 360
Pret/Punret 1.35 0.50
With Nails
Specimen SP14 SP15 SP16
Punret [kN] 302 498 632
Pret [kN] 351 498 623
Pret/Punret 1.16 1.00 0.99
is important to highlight cases such as SP8 and SP16
with the same size and yet different strengths due to dif-
ferent level of deterioration and defects in the wood.
The maximum forces attained in the retrofitted tests
were similar to or greater than those obtained from the
unretrofitted specimens regardless of the retrofitting de-
tails, with the exceptions of specimen SP5 and SP13. The
strength of the retrofitted specimen SP5 was found to be
91% of that of the unretrofitted specimen. This slight
reduction in strength could possibly be due to imper-
fectly wrapped FRP sheets since the surface of the
specimen was quite uneven. Note that the specimens with
mortar shell, which made the surface of the specimens
more evenly round, showed no reduction in strength. For
the case of SP13, due to the weak out of plane resistance
of CFRP jacket, an early crushing of the filling material
(mortar) in the wedge cut between the two halves of the
timber pile specimen was observed. This weakness was
produced by the extremely small thickness of the CFRP
jacket. Further, the results shown in the table prove that
using steel drift pins or nails results in similar perform-
ance, since the average retunret
PP ratio for these cases
Copyright © 2012 SciRes. OJCE
were 1.06 and 1.05, respectively.
To understand better how the behavior of retrofitted
specimens compares to that of unretrofitted specimens,
some selected force-displacement relationships for dif-
ferent retrofit details are presented in Figure 14. In this
figure, the displacements are obtained from the internal
LVDT of the compression machine, i.e. the displace-
ments at the point of application of the compression load,
76 mm away from the centroid of the pile circular section.
Early in the loading history, up to approximately 100 kN
both unretrofitted and retrofitted specimens had a softer
behavior, in all cases, likely due to the flexibility of the
loading frame. The retrofitted test was subsequently
slightly softer than the unretrofitted test for cases (a) and
(d) in Figure 14. This behavior is likely due to the slight
damage that was exerted on the specimen when it was
first tested while it was still unretrofitted. Moreover, it
was found that the use of mortar shell together with FRP
sheets (Figure 14(b)) helped in restoring the strength and
stiffness of the pile. Hence, using mortar shell could be
recommended in the cases of piles with higher levels of
damage. In the case where a mortar-filled wedge is in-
troduced prior to wrapping the specimen with FRP, since
the elastic modulus of mortar was higher than the elastic
modulus of the wood, the retrofitted pile showed higher
stiffness than that of the unretrofitted pile at the initial
stage (Figure 14(c)). However, as the testing progressed,
the specimen became softer due to the weak out of plane
resistance of CFRP jacket, as mentioned earlier, which
provided very little confinement for the filling mortar
resulting in its early crushing. From Figure 14(d), it is
found that using nails or drift pins does not impact sig-
nificantly the capacity or stiffness of the retrofitted piles.
In an elastic design of the pile cap and piles, forces are
distributed from the pile cap to the piles according to the
stiffness of each pile, since the pile cap can be regarded
as a rigid body. Therefore, it was important to ensure that,
although the strength of the retrofitted specimens was
satisfactory, their stiffness is not significantly impacted.
To eliminate the effect of early hardening which was
highlighted earlier in the discussion, the stiffness was
calculated after the specimen reaches a stable linear be-
havior. It was found by inspection that this behavior oc-
curs approximately between one third and two thirds of
the maximum force. Therefore, the stiffness results pre-
sented in this study were calculated using the middle
portion of force-displacement curves.
(a) Without mortar shell (b) With mortar shell
(c) With wedge (d) With nails
Figure 14. Force-displacement relationship for specimens with different retrofit conditions.
Copyright © 2012 SciRes. OJCE
A comparison between the unretrofitted and retrofitted
stiffness of each specimen is presented in Figure 15. The
results of CFRP retrofitted specimens are excluded in
Figure 15, since SP13 showed a different type of failure
mechanism and also the results produced higher standard
deviation when added to the GFRP test results; hence it
was more conservative to eliminate them. SP11 was not
presented either, since it experienced relatively extensive
damage during the first (unretrofitted) test. The retrofit-
ted specimens SP1, SP2 and SP8 show higher stiffness
than that of the unretrofitted specimens since damage
experienced in the specimens while conducting the un-
retrofitted test was extremely minor. By using mortar
shell, the stiffness of the retrofitted specimens was fully
restored, and the stiffness of specimens SP8R and SP10R
was found to be 112% and 99% of the unretrofitted
specimens, respectively. The stiffness results presented
in Figure 15 showed a mean stiffness reduction of 5%
and median stiffness reduction of 14% for the retrofitted
specimens compared with the unretrofitted specimens.
7. Design Recommendation
Figure 16 shows a graph of the ratios of the maximum
retrofitted forces to the maximum unretrofitted forces
with respect to the GFRP volumetric ratio used for the
retrofitted specimens. A linear regression analysis was
conducted and the result is shown as a solid straight line
in the figure. The two dashed lines drawn in the figure
represent 2
, where,
is the mean and
the standard deviation of the data points, respectively. As
shown in the figure, only two data points lie below the
lower dashed line; hence it was deemed appropriate to
base the design on the 2
value. The design point is
considered when the
line intersects the hori-
zontal line corresponding to a retunret ratio equal to
1.0. Therefore, the design volumetric ratio of GFRP is
found to be 0.068. Note that only the tests with GFRP
were used to calculate the results shown in Figure 16. It
is important to note also that the average ratio of the ret-
rofitted to the unretrofitted stiffnesses of the five speci-
mens whose volumetric ratios is higher than the proposed
value of 0.068 is 1.08. This confirms that while similar
strength can be expected by using the proposed volumet-
ric ratio, the stiffness of the retrofitted specimens will not
be jeopardized.
It is important to recall that the calculated volumetric
ratio corresponds to the GFRP type used in this study.
However, it can be generalized to accommodate other
FRP types with different modulus of elasticity, EFRP. The
equivalent design FRP volumetric ratio (
RP eq
) can be
expressed as follows:
is the design volumetric ratio of FRP used
in this study with
RP equal to 12400 MPa. The number
of layers n, corresponding to a layer thickness equal to t,
needed to retrofit the timber pile post connection can be
calculated as:
2FRP eq
 (9)
Finally, Equations (8) and (9) can be merged to form
the following equation:
As stated above, based on the experimental tests and
after regression analysis, the optimal volumetric ratio
is found to be 0.068. Therefore, Equation (10) can
be simplified to:
Figure 15. Stiffness of unretrofitted and retrofitted speci-
Figure 16. Regression analysis of the FRP volumetric ratio
vs. Pret/Punret data points.
Copyright © 2012 SciRes. OJCE
where the unit of EFRP is MPa.
8. Conclusions
This study focused on evaluating the feasibility of using
FRP retrofit technique to restore the flexural capacity of
deteriorated bridge timber piles under eccentric loading.
A total of 16 pile specimens were used in the study. Each
specimen was tested before and after retrofitting with
GFRP or CFRP sheets. To assess the impact of realistic
field conditions, different details of the FRP retrofit
technique were investigated including applying FRP with
and without mortar shell, introducing mortar-filled wedge
in the tested specimen prior to wrapping it with FRP to
mimic the effect of severe damage, and posting the piles
with nails instead of steel drift pins. The test results
showed acceptable structural behavior for the specimens
retrofitted with GFRP sheets, regardless of the retrofit
details adopted in the tests. On average, specimens retro-
fitted with GFRPs showed strength 10% greater than that
of unretrofitted specimens. The behavior of CFRP sheets
however was less satisfactory due to the small thickness
of the CFRP shell used as a result of the high strength of
CFRP compared to GFRP. This resulted in excessive out
of plane deformation of CFRP shell under large com-
pressive stress. It was found from the study that using
mortar shell along with FRP sheets helps in enhancing
the stiffness of the retrofitted pile; hence is recommended
for piles with higher levels of deterioration. The average
mean and median reduction of stiffness of the retrofitted
specimens was 5% and 14% of the unretrofitted speci-
mens, respectively. Linear regression analysis was con-
ducted on the test data, and a design value for the FRP
volumetric ratio of 6.8% was determined for the GFRP
type used in this study. This value was utilized along
with the elastic design theory to propose a more generic
formula for the design of retrofit for bridge timber piles
using any FRP type.
9. Acknowledgements
The authors acknowledge the support for this study pro-
vided by the Illinois Center for Transportation (ICT) and
the Illinois Department of Transportation (IDOT) under
project No. R27-82.
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