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Fiber‐Reinforced Plastic (FRP) Wraps for Next Generation Sustainable and Cost‐Effective Rehabilitation of Coastal Transportation Infrastructure in the Mid‐Atlantic Region
Date: March 2018
Wael, Zatar, PhD, Professor, Marshall University Hai, Nguyen, PhD, Research Engineer, Marshall University Osman, Ozbulut, PhD, Assistant Professor, University of Virginia Prepared by: Marshall University Research Corporation 1 John Marshall Drive Huntington, WV 25755
Prepared for: Virginia Center for Transportation Innovation and Research 530 Edgemont Road Charlottesville, VA 22903
FINAL REPORT
I
1. Report No.
2. Government Accession No.
3. Recipient’s Catalog No.
4. Title and Subtitle
Fiber‐Reinforced Plastic (FRP) Wraps for Next Generation
Sustainable and Cost‐Effective Rehabilitation of Coastal
Transportation Infrastructure in the Mid‐Atlantic Region
5. Report Date
March 2018
6. Performing Organization Code
7. Author(s)
Wael Zatar, Hai Nguyen, and Osman Ozbulut
8. Performing Organization Report No.
9. Performing Organization Name and Address
Marshall University Research Corporation
10. Work Unit No. (TRAIS
11. Contract or Grant No.
DTRT13-G-UTC33
12. Sponsoring Agency Name and Address
US Department of Transportation Office of the Secretary-Research UTC Program, RDT-30 1200 New Jersey Ave., SE Washington, DC 20590
13. Type of Report and Period Covered
Final 03/01/16 – 08/31/17
14. Sponsoring Agency Code
15. Supplementary Notes
16. Abstract
Transportation infrastructure in mid-Atlantic region (including the District of Columbia, Delaware, Maryland, Pennsylvania, Virginia, and West Virginia), particularly concrete highway bridges, are gradually exposed to the deleterious effects of environmental attacks, leading to environmental degradation of the concrete materials. This is due to, for example, carbonation and chloride contamination that eventually break the alkali barrier in the cement matrix, and the steel reinforcement in the concrete becomes susceptible to corrosion. As a consequence, the concrete may deteriorate at the reinforcement level, leading to cracking and spalling of the concrete owing to volume increase of the steel reinforcement. Such degradation is exacerbated by the application of de-icing salts on highway bridges, and the freeze-thaw and dry-wet cyclic exposures causing accelerated ageing of the structure over time. Concrete deterioration in the United States and worldwide has motivated the development of new and innovative materials and methods for structural rehabilitation, since replacement of structures would be very costly and nearly prohibited. One solution to overcome steel corrosion in concrete for new construction is to use Fiber-Reinforced Polymer (FRP) materials for internal reinforcements instead of steel. More significant is the beneficial application of FRP for structural rehabilitation of deteriorated concrete structures. FRP composite materials in the form of fabrics, laminates, and bars have been externally bonded to concrete structures to increase structural capacity and provide longer service-life. The application of this technology in practice has been highly successful. The main goal of this project is to present a comprehensive review and technical/economical effectiveness of externally-bonded FRP composites for repair and retrofit of highway infrastructure, and particularly concrete bridges.
17. Key Words
Mid-Atlantic region; Concrete bridges; Deterioration; Structural rehabilitation; FRP composites; Externally bonded
18. Distribution Statement
No restrictions. This document is available from the National Technical Information Service, Springfield, VA 22161
19. Security Classif. (of this report)
Unclassified
20. Security Classif. (of this page)
Unclassified
21. No. of Pages
128
22. Price
II
EXECUTIVE SUMMARY
A major impediment for the implementation of FRP in transportation infrastructure is the lack
of effective references for training and technology transfer in practice, because the information
is usually disparate, often confusing, and even contradictory. Therefore, this project offers the
opportunity for developing a concise yet complete reference report that can serve as a
“practical” educational tool, and facilitate the evaluation and implementation of externally
bonded FRP repair by State DOT personnel. An extensive and critical review of FRP repairs
for highway structures/elements are performed. Applicable and sound NDT/NDE process and
practices for inspecting FRP external wraps are introduced. Existing reports, national
guidelines, specifications for design and construction of externally bonded FRP are reviewed.
The reviews cover all the aspects of FRP research pertaining to repairing, reinforcing, or
strengthening by external wrap and near-surface mounting. Aspects of materials, design,
construction, durability, and maintenance are covered. National guidelines pertaining to FRP-
retrofit including ACI and NCHRP are evaluated. Cost of few FRP-wrap projects by WVDOT
and other state DOTs are addressed. The research team works closely with WVDOH and
VDOT members managing the databases, inventory, inspection reports, digital pictures, load
postings, etc. Details of few FRP-retrofitted projects in West Virginia and Virginia are
provided. Bridge location, purpose of the FRP wrap, retrofit details are documented. Overall
conditions of all highway bridges in the state of Virginia and West Virginia are reported. These
data are extracted from the latest National Bridge Inventory by U.S. Department of
Transportation, Federal Highway Administration. FRP design spreadsheet for flexural
strengthening of RC T-beams is presented. The developed spreadsheet employs the NCHRP
Report 655 “Recommended Guide Specification for the Design of Externally Bonded FRP
Systems for Repair and Strengthening of Concrete Bridge Elements” and the AASHTO LRFD
Bridge Design Specifications, 7th Edition (AASHTO 2014).
III
ACKNOWLEDGEMENTS
This project was sponsored by the United States Department of Transportation’s University
Transportation Centers (USDOT UTC) Program. The principal investigator of the project, Wael Zatar,
would like to acknowledge the supports provided by the USDOT UTC and the Mid-Atlantic
Transportation Sustainability University Transportation Center (MATS UTC). The principal
investigator of the project would also like to acknowledge the great support provided by Donald
Williams of the West Virginia Department of Transportation.
IV
DISCLAIMER STATEMENT
The contents of this report reflect the views of the authors, who are responsible for the facts and the
accuracy of the information presented herein. This document is disseminated under the sponsorship of
the U.S. Department of Transportation’s University Transportation Centers Program, in the interest of
information exchange. The U.S. Government assumes no liability for the contents or use thereof.
1
Table of Contents
EXECUTIVE SUMMARY ...................................................................................... II
ACKNOWLEDGEMENTS .................................................................................... III
DISCLAIMER STATEMENT ............................................................................... IV
TABLE OF CONTENTS ........................................................................................... 1
CHAPTER 1 – INTRODUCTION ............................................................................ 3
1.1 Introduction ............................................................................................................................4
1.2 Research statement and objectives .........................................................................................5
CHAPTER 2 – LITERATURE REVIEW OF STRUCTURAL RETROFITTING
USING FIBER REINFORCED POLYMERS .......................................................... 7
2.1. Introduction ...........................................................................................................................8
2.2. Mechanical properties and manufacturing processes of FRP composites ............................8
2.3. Upgrading concrete structures using FRP composites ........................................................11
2.3.1. Flexural strengthening .............................................................................................12
2.3.2. Shear strengthening ..................................................................................................30
2.3.3. Column strengthening and confinements.................................................................32
2.4. Upgrading metallic structures .............................................................................................37
2.5. Upgrading timber and masonry structures ..........................................................................38
2.6. NDT/NDE process for inspecting FRP wraps .....................................................................40
2.7. National guidelines pertaining to FRP-retrofit ....................................................................41
2.8. Cost of FRP wraps ...............................................................................................................55
2
CHAPTER 3 –FRP-RETROFITTED BRIDGE PROJECTS BY WVDOT &
VDOT ....................................................................................................................... 57
3.1. FRP-retrofitted bridge projects by WVDOT .......................................................................60
3.2. FRP-retrofitted bridge projects by VDOT .........................................................................70
CHAPTER 4 –DESIGN SPREADSHEET FOR FLEXURAL STRENGTHENING
OF RC BEAMS USING FRP COMPOSITES ........................................................ 72
4.1. Evaluation of concrete structures prior to rehabilitation .....................................................73
4.2. Sufficiency rating and overall bridge conditions in Virginia and West Virginia ................74
4.3. Design of flexural strengthening of RC T-beams using FRP ..............................................77
CHAPTER 5 – PROJECT SUMMARY AND CHALLENGES ........................... 112
5.1. Project summary and challenges .......................................................................................113
REFERENCES ....................................................................................................... 113
3
Chapter 1
Introduction
4
1.1 Introduction
In the United States, about 26% of the highway bridges are in need of repair or replacement, and
a large number of these deficient bridges are reinforced or pre-stressed concrete structures. The
cost of the US infrastructure rehabilitation is estimated at over 1.5 trillion dollars over the next five
years, with corrosion deterioration costs due to deicing and sea salt effects estimated at $150
billion. The United States Congress has recently approved a multi-year, $305 billion highway,
transit and railway authorization bill to provide much-needed funds for state DOTs to fix
deteriorated and deficient bridge infrastructure. Any rational decision regarding maintenance,
repair, or replacement of the deteriorated members should take into account the member’s
condition, the extent of deterioration, the expected remaining service life, and the impact of
alternative maintenance and repair options on the service life of the members. In making these
decisions, the state DOTs across the country use visual inspection techniques, among others, for
monitoring the extent of cracking and damage/deterioration of the transportation infrastructure.
The use of FRP wraps to reinforce or strengthen concrete elements is a practice that has been
addressed with some success nationally and internationally. In West Virginia, over the past 20
years, various projects have benefitted mainly from having FRP as an internal reinforcement.
Applications ranging from decks to strengthening of bridge components have been accomplished,
mostly with success. As funds are anticipated to be limited in the future, the ability of designers to
look at alternative means of replacing or strengthening structures is important. It is anticipated that
there are various structures in the inventories of both West Virginia and Virginia that could benefit
by having members strengthened or stabilized with the use of FRP wraps. In the past, the process
for doing the work was on a case by case basis with independent plans and specifications for each
project developed. Often times propriety materials were used for the repairs. The goal of this
project is to inventory the processes that are available in repairing or strengthening concrete bridge
elements. The possible pool of candidates that would benefit from the strengthening/rehabilitation
will also be evaluated. The ultimate goal is to determine which structures are available for repair,
how many structures could benefit from this process, and to determine the next phase of developing
specifications and standard details for the repairing or strengthening.
5
1.2 Research statements and objectives
Transportation infrastructure, and particularly concrete highway bridges are exposed over time
to the deleterious effects of environmental attacks, leading to environmental degradation of the
material due for example to carbonation and chloride contamination that eventually break the
alkali barrier in the cement matrix, and the steel reinforcement in the concrete becomes
susceptible to corrosion. Consequently, the concrete may delaminate at the reinforcement level,
leading to cracking and spalling of the concrete due to volume increase of the steel reinforcement.
Such degradation is exacerbated by the application of deicing salts on highway bridges, and the
freeze-thaw and dry-wet cyclic exposures causing accelerated ageing of the structure over time.
In the United States, about 26% of the highway bridges are in need of repair or replacement, and
a large number of these deficient bridges are reinforced or pre-stressed concrete structures. The
cost of the United States infrastructure rehabilitation is estimated at over 1.5 trillion dollars over
the next five years, with corrosion deterioration costs due to deicing and sea salt effects estimated
at $150 billion. The United States Congress has recently approved a multi-year, $305 billion
highway, transit and railway authorization bill to provide much-needed funds for state DOTs to
fix deteriorated and deficient transportation infrastructure.
The West Virginia Division of Highways is in charge of maintaining nearly 7,228 bridges
over 10 districts, with concrete bridges representing about 46% of the inventory (about 17% cast-
in-place and 29% pre-cast concrete); in addition, there are about 350 concrete culverts. The
transportation infrastructure system in Virginia is one of the largest in the nation, with very
similar salt-related deterioration issues resulting from its winter snow and ice control. Aging
transportation infrastructure in Virginia are in a worse situation than West Virginia because of
the continuous exposure to the coastal environment, with chloride penetration being a huge factor
in accelerating the infrastructure deterioration. Concrete deterioration in the United States and
worldwide has motivated the development of new and innovative materials and methods for
structural rehabilitation, since replacement of structures would be very costly and nearly
prohibited. One solution to overcome steel corrosion in concrete for new construction is to use
Fiber-Reinforced Polymer or Plastic (FRP) materials for internal reinforcements instead of steel.
More significant is the beneficial application of FRP for structural rehabilitation of deteriorated
concrete structures. FRP composite materials in the form of fabrics, laminates, and bars have
been externally bonded to concrete structures to increase structural capacity and provide longer
6
service-life. The application of this technology in practice has been highly successful. The main
goal of this proposal is to investigate the technical and economical effectiveness of externally-
bonded FRP composites for repair and retrofit of highway infrastructure, and particularly
concrete bridges.
The geographical location of MATS States, with environmental and coastal-induced
deficiencies, and the need for salt spraying during the wintertime produce the greatest challenges
for bridge infrastructure maintenance, sustainability and asset management efforts. The primary
objective of this project is to provide a comprehensive review and technical/economical
effectiveness of FRP composites for sustainable retrofitting and rehabilitation of deficient
concrete in the transportation infrastructure systems of both West Virginia and Virginia. While
complementing the ongoing efforts that have already been funded through the West Virginia State
Planning and Research (SP&R) funds and approved by the FHWA, this project will focus on
additional parameters that could play a major role on assessing the efficacy of FRP wraps for
coastal infrastructure.
7
Chapter 2
Literature Review of
Structural Retrofitting
Using Fiber Reinforced
Polymers
8
2.1. Introduction
According to Portland Cement Association (PCA), concrete can deteriorate for variety of reasons
such as carbonation (occurs when carbon dioxide from the air penetrates the concrete and reacts
with hydroxides, such as calcium hydroxide, to form carbonates), free-thaw deterioration,
chemical attacks (e.g. by acids, alkalis, and sulfate), alkali-aggregate reactivity, abrasion/erosion,
unusually high temperatures (e.g. fire, heat), plastic and drying shrinkage cracking, thermal
stresses, overloading and impact, loss of support, and surface defects (e.g. surface air voids,
delaminations, etc.). Of these reasons, corrosion of reinforcing steel is the leading cause of
deterioration in concrete. The corrosion of reinforcing steel is especially critical in aggressive
marine environments and in cold regions, where de-icing salts are used. Of the 614,387 existing
bridges in the United States, 9.1% of the nation’s bridges (i.e. 56,007 bridges) were structurally
deficient in 2016 and about 40% are 50 years or older. The average age of America’s bridges keeps
going up and many of them are approaching the end of their design life. The most recent estimate
puts the nation’s backlog of bridge rehabilitation needs at $123 billion (ASCE 2017). It is therefore
important to find material, technology, and economical solutions to effectively address the
deficient bridges. The need for economically retrofitting of structures in the US and worldwide
has led to numerous research in FRP strengthening.
The practice of bonding carbon/epoxy composites to reinforced concrete beams to increase
their flexural capacity was first reported in the mid-1980s, when Germany and Switzerland
replaced steel with FRP plates (Hollaway 2011). Externally-bonded FRP composites to strengthen
RC and PC members in flexural and shear are viewed as an effective method to enhance structures’
strength/stiffness. This chapter presents mechanical properties and manufacturing process of some
important FRP composites that are successfully used for structural strengthening and seismic
retrofitting. An extensive and state-of-the-art review on FRP strengthening/retrofitting is provided.
Non-destructive evaluation technique for inspecting FRP wraps and national guidelines pertaining
to FRP-retrofit are introduced.
2.2. Mechanical properties and manufacturing process of FRP composites
Carbon, glass, and aramid (a.k.a. Kevlar, DuPont’s registered trademark for a para-aramid
synthetic fiber) fibers in forms of sheets (fabrics) or laminates (plates, strips) are commonly used
9
to strengthen RC and PC structural members in flexure and shear. Typical mechanical properties
of these fibers are listed in Table 2.1.
Table 2.1 Typical Mechanical Properties of Carbon, Glass, and Aramid Fibers (Hyer 2009;
Nguyen et al. 2017a)
Fiber Density
(g/cm3)
Diameter
(m)
Elongation
(%)
Tensile
strength
(MPa)
Young’s
modulus
(GPa)
PAN-based
Carbon (IM) 1.78-1.82 8-9 1.0 2,410-2,930 228-276
PAN-based
Carbon (HM) 1.67-1.9 7-10 0.5 2,070-2,900 331-400
PAN-based
Carbon
(UHM)
1.86 7-10 0.3-0.4 1,720 517
E-glass 2.54 8-14 1.8-3.2 3,450 72.4
S-glass 2.49 10 5.7 4,590 85.5
Aramid
(Kevlar-29) 1.44 12 3-4 2,760 62
Aramid
(Kevlar-49) 1.48 12 2.2-2.8 2,800-3,792 131
Note: PAN = Polyacrylonitrile; IM = Intermediate Modulus; HM = High Modulus; UHM = Ultra-
High Modulus.
Glass fibers are by far most commonly used artificial fibers on earth. They possess many good
mechanical properties such as high specific strength/stiffness, low cost, low density, good fire and
chemical resistance, and good electric insulation (Nguyen et al. 2017b). However, under certain
conditions of exposure, glass fibers are sensitive to alkaline environments and moisture attack and,
in addition, creep affects glass fibers more than any other types of fibers (Hollaway 2011).
Carbon fiber reinforced polymer (CFRP) composites were developed during the 1960s for
specialized applications. Unlike glass fiber, carbon fiber is an electrical conductor and hence
10
galvanic corrosion could take place if fibers are placed in direct contact with metals. However, this
effect is irrelevant when CFRP is bonded to RC members. In general, CFRP composites exhibit
excellent fatigue and creep properties (Hollaway 2011). High performance carbon fibers can be
achieved by using the precursor polyacrylonitrile (PAN).
Aramid fibers were first developed in 1965 and likewise do not creep or fatigue under load.
They have anisotropic mechanical properties, with higher strength and modulus of elasticity values
in their longitudinal direction compared to their transverse direction (Hollaway 2011). In general,
Kevlar has a unique combination of high strength, high modulus, toughness and thermal stability.
Vinylesters and epoxies are two major polymers used with glass, carbon, and aramid fibers to
form FRP composites. There are three main manufacturing methods of FRP composites for
retrofitting RC and PC structural systems including (1) manual process (e.g wet lay-up); (2) semi-
automated process (e.g. pre-impregnated fiber (prepreg) molding and filament winding); and (3)
automated process (e.g. pultrusion). Each method will have an effect upon the quality, performance
and therefore characteristics of the final composite. The automated fabrication methods have a
high degree of production control, composite compaction, and complete cure compared to the
manual fabricated techniques. Therefore, the automated process will have higher values of strength
and stiffness compared to those of other techniques (Hollaway 2011). An overview of all
techniques/processes for manufacturing FRP composites is illustrated in Figure 2.1.
11
Note: SMC = Sheet Molding Compound; SRIM = Structural Reaction Injection Molding;
BMC = Bulk Molding Compound; RTM = Resin Transfer Molding; SCRIMP = Seeman
Composite Resin Infusion Molding Process
Figure 2.1 Manufacturing processes for polymer matrix composites (Mazumdar 2001
and Nguyen et al. 2017)
2.3. Upgrading concrete structures using FRP composites
Reinforced concrete members have been traditionally upgraded using externally epoxy-bonded
steel plates. This technique is simple, efficient, and cost-effective and but it has the following
drawbacks: the corrosion of steel plates leads to deterioration of the bond at the steel-concrete
interface; manipulating the heavy steel plates is challenging at the construction site; scaffolding is
required; and in the case of flexural strengthening of long elements, lengths of steel plates are
limited by sizes and load-carrying capacity of delivery trucks. Alternatively, the steel plates can
be replaced by FRP sheets or strips (Bakis et al. 2002). Another method for strengthening RC
columns is steel jacketing. This method is proved to effectively enhance shear strength and
Composites
processing
Composites
processing
Thermoset
composites
Thermoset
compositesThermoplastic
composites
Thermoplastic
composites
Short-fiber
composites
Short-fiber
compositesContinuous-fiber
composites
Continuous-fiber
composites
· SMC
· SRIM
· BMC
· Spray-up
· Injection molding
· SMC
· SRIM
· BMC
· Spray-up
· Injection molding
· Filament winding
· Pultrusion
· RTM
· Hand lay-up
· Autoclave
· Roll wrapping
· SCRIMP
· Bladder molding
· Filament winding
· Pultrusion
· RTM
· Hand lay-up
· Autoclave
· Roll wrapping
· SCRIMP
· Bladder molding
Short-fiber
composites
Short-fiber
compositesContinuous-fiber
composites
Continuous-fiber
composites
· Injection molding
· Blow molding
· Injection molding
· Blow molding· Thermoforming
· Tape winding
· Compression
molding
· Autoclave
· Diaphragm
forming
· Thermoforming
· Tape winding
· Compression
molding
· Autoclave
· Diaphragm
forming
12
ductility of columns (Priestley et al. 1994a-b). However, it increases the dead loads and the cross-
sectional dimensions of the structure, resulting in a potentially undesirable stiffness increase. As
an alternative, concrete elements in existing structures can be externally reinforced by high-
strength FRP composites with desired length, width, and thickness of FRP wraps. Furthermore,
FRP wrapping may be tailored to meet specific structural requirements by adjusting the placement
of fibers in various directions (Bakis et al. 2002).
2.3.1. Flexural Strengthening
Externally bonded FRP
The flexural strength of a reinforced concrete beam can be enhanced by a FRP plate bonded to
the tension face or the soffit of simply supported beams (Smith and Teng 2002). Over the past
30 years, researches on FRP-plate bonding for strengthening of RC beams has been well
established. Very first study on strengthening RC beams using CFRP strips was conducted at
the EMPA (the German acronym for Swiss Federal Laboratories for Materials Testing and
Research) in mid-1980s. The world first application of the CFRP strips was the Ibach bridge,
Switzerland in 1991. The box beams of this bridge was successfully strengthened by the CFRP
strips of 5 m length (Meier 2000). The externally-bonded FRP technique for flexural
strengthening of beams has been widely applied in practice in recent years as it offers many
advantages such as good corrosion resistance, ease of site handling, and minimum increases in
structural weight and size. Numerous researches have been conducted to evaluate the behavior
and strength of FRP-strengthened RC beams (Norris et al. 1997, Saadatmanesh and Malek
1998, Buyukozturk and Hearing 1998, Teng et al. 2002, Hollaway and Teng 2008…).
Huang et al. (2016) studied the effect of flax fabric reinforced polymer(FFRP) on beam
strengthening by externally bonded reinforcement. Test variables included FFRP thickness, the
steel reinforcement ratio and the pre-cracking of RC beams. Based on the experimental results,
conclusions can be drawn as follows:
· FFRP strengthening increases the ultimate load capacity. Beams with more FFRP layers
have higher ultimate load. The increase in load carrying capacity for RC beams with a low
steel reinforcement ratio due to FFRP strengthening is larger than that of the beams with a
high reinforcement ratio.
13
· FFRP strengthening increases the deflection of the RC specimens remarkably. The increase
in deflection is more propounded for beams with a lower steel reinforcement ratio. In
addition, beams with more FFRP layers have larger deflection.
· ductility and energy absorption capacity increase significantly due to the FFRP
strengthening. However, the increase in ductility and energy absorption of RC beams with
low steel reinforcement ratio is larger than that in high-reinforcement ratio beams.
· The pre-cracking slightly increases the deflection at failure, ductility and ultimate FRP
strain of the strengthened beams, but has no obvious effect on the failure mode, ultimate
load capacity and energy absorption of these beams.
· The FFRP strengthened beams show the same failure mode, i.e. initial steel yielding,
followed by brittle rupture of the FFRP laminate. For all the strengthened beams, no
debonding of FFRP plates from concrete is observed, indicating a good compatibility
between the FFRP plates and the RC beam due to the use of U-shaped flax FRP bands at
the ends of the beams.
· Compared with the control RC beams, the FFRP strengthened beams show more and wider
cracks at failure.
· The comparison with RC beams having similar dimension but externally bonded by GFRP,
CFRP and steel plates indicates that the enhancement in ultimate lateral load carrying
capacity due to natural FFRP plates is close to or comparable to the GFRP, CFRP and steel
plate strengthened beams, although the tensile strength and modulus of FFRP composites
are significantly lower than those of CFRP, GFRP and steel.
Yasir (2016) studied the flexural behavior of beams strengthened with prestressed CFRP
strands and investigated the new technique involved anchoring the CFRP strands at the ends of
the concrete beams using a new "steel tube" anchorage system. Based on the experimental
results, conclusions can be drawn as follows:
· The expansive grout (Bustar), used as the filler material for the anchorage system, provided
adequate pressure which maintained the CFRP strand inside the steel tube until the full
14
tensile strength of CFRP was achieved. This resulted in a successful steel tube anchorage
system used in one beam.
· Both lengths (15 in. and 12 in.) of the CFRP anchors achieved more than the guaranteed
tensile capacity of the CFRP strand. However, the 15-in. long anchorage system could
sustain 130% the guaranteed capacity with a load-slip stiffness much higher than that
achieved by 12-in.-long anchors.
· The confinement provided by lateral reinforcement significantly affected the CFRP strand-
to-concrete bond characteristics.
· The transfer lengths measured at the live and dead ends were found to be almost the same.
· The average bond stress at transfer increased when the transfer length decreased as a result
of adding more lateral reinforcement.
· Adding the CFRP steel tube anchorage system at the ends increased the average bond stress
at transfer by about 60% and decreased the transfer length by about 36%
· CFRP members need to have adequate lateral confinement in order to prevent early bond
failures.
· Based on the experimental results, the average total losses in prestressing force from the
jacking up to the flexural test day (typically an average of 25 days) can be estimated as
7.3% for CFRP strands.
· The devised technique of using steel tube anchorage system at the ends of CFRP strands
prevented the end slippage. In this study, using steel tube anchorage system improved the
member flexural capacity by 33%.
· Comparing CFRP beams with beams prestressed with steel strands, the CFRP beams
showed higher strength but less ductility when both beams had the same cross-sectional
area, prestressing force, span length, and designed for the same service load.
Spadea et al. (2015) studied the structural behavior of concrete beams with EBR FRP
systems. Strength increase, ductility, and the ability to dissipate the internal strain energy were
analyzed by the experiment and conclusions are drawn as follows:
· All the unplated beams failed in the flexural conventional mode, after extensive yielding
of tension steel followed by crushing of concrete in the compression zone.
15
· The beams with only the CFRP laminate bonded to the tension face, without any anchorage
system carried higher loads, compared to the control beams, but all failed in a brittle
manner with the load capacity dropping suddenly by explosive debonding of FRP plates.
· The end and other supplementary anchorage devices favored a more ductile failure and an
increase of the load carrying capacity.
· In all cases the failure occurred in a sudden way, so there was not the possibility to measure
the exactly value of the increase of strength.
Hawileh et al. (2014) studied the behavior of Reinforced Concrete (RC) beams strengthened
in flexure by means of different combinations of externally bonded hybrid Glass and Carbon
Fiber Reinforced Polymer (GFRP/CFRP) sheets. Based on the experimental results, conclusions
can be drawn as follows:
· The combination of the Carbon/Glass sheets decided the increment of load capacity of the
strengthened beams.
· The ductility at failure for the beams strengthened with glass and hybrid sheets is higher
than that with a single carbon sheet. The ductility at failure of the beam strengthened with
a single glass sheet is the highest among all the strengthened beams.
· The use of hybrid systems combines the lower stiffness of the GFRP sheets with the high
strength of the CFRP sheets to result in a material, which provides an improved strength
and ductility in beam behavior.
· The failure modes for the tested beam specimens showed different local and global failure
characteristics including concrete crushing, flexural cracks, debonding, and delamination.
Burke et al. (2013) studied the performance of loaded externally bonded and NSM FRP
flexural strengthening systems for reinforced concrete structures at temperatures that they might
experience if well-insulated and exposed to a standard fire scenario and the relationship
between Tg, load, exposure temperature, and structural effectiveness for both externally bonded
and NSM FRP strengthening systems for concrete structures and testified whether the high
temperature performance of the NSM FRP strengthening system could be improved by using a
cementitious, rather than an epoxy adhesive. Conclusions are drawn as follows:
16
· At ambient temperature, the Epoxy adhesive provides superior bond performance as
compared with the cementitious grout adhesive for NSM FRP strengthening systems for
concrete.
· If well insulated against the thermal effects of fire, it may be possible for NSM FRP
strengthening systems to achieve structural fire endurance ratings of several hours.
· the performance at high temperature of NSM FRP strengthening systems can be improved
by using a cementitious grout adhesive rather than an ambient temperature cure epoxy.
· The externally bonded FRP system had better performance at elevated temperature than
other systems tested.
Fanning and Kelly (2001) investigated the flexural behavior of ten reinforced concrete
beams strengthened with different plate configurations. They found that the external bonding
of CFRP plate offers an extremely effective means of strengthening RC beams in flexure. The
results showed a 40% increase in stiffness and a corresponding 70% increase in ultimate
strength for reaction anchored plates. The effectiveness of the external plates in strengthening
reduced as the plate lengths were shortened. A strain compatibility and force equilibrium
method of analysis, coupled with an empirical rule derived from the test data, was
demonstrated to be effective in predicting the ultimate response of simply supported beams in
bending with and without end plate anchorages and irrespective of plate length.
Baghiee et al. (2009) conducted a series of dynamic and static tests on beams strengthened
with CFRP sheets. Vibration-based monitoring techniques were used as they are proven useful
in identifying the changes of structural properties due to damage and strengthening. Six of the
specimens were strengthened with externally bonded CFRP sheets when the load reached
approximately half of the predicted failure load. The results showed that the frequencies were
affected by damage and strengthening, but their changes were not influenced by damage
locations. The frequencies decreased with increasing damage, however, they were also affected
by environmental conditions (e.g. ambient temperature). The Modal Assurance Criterion
(MAC) was found to be subjected to very small change by damage or strengthening (according
to Allemang (2003), MAC were originated from the need for a quality assurance indicator for
experimental modal vectors that were estimated from measured frequency response functions).
The MAC values can reveal the changes of overall stiffness of the beams during the load steps.
17
The change of stiffness at each degree of freedom of beams evaluated by Coordinate MAC
(COMAC) and modal curvatures indicated that the damage identification of the beam
specimens was best described by modal curvature method.
Grace et al. (1999) studied the behavior of reinforced concrete beams strengthened with
various types of fiber reinforced polymer (FRP) laminates. Fourteen simply-supported beams
with rectangular cross sections were strengthened and tested. Each beam was initially loaded
above its cracking load. The cracked beams were strengthened with FRP laminates and then
tested until complete failure. Five available strengthening systems of various types of
carbon/glass fiber reinforced polymer (CFRP/GFRP) strengthening materials were used. These
materials included two types of CFRP sheets, bi-directional and unidirectional GFRP sheets,
and CFRP plates.
Triantafillou and Plevris (1992) established a systematic analysis procedure for the short-
term flexural behavior of FRP-strengthened members. FRP-strengthened concrete beams can
fail in several ways when loaded in bending. The following collapse mechanisms were
identified and analyzed in this study: steel yield-FRP rupture, steel yield-concrete crushing,
compressive failure, and debonding. They obtained equations to describe each failure
mechanism using the strain compatibility method, concepts of fracture mechanics, and a simple
model for the FRP peeling-off debonding mechanism due to the development of shear cracks.
They recommended that the additional issues, such as behavior under sustained loading,
fatigue, thermal cycling, humidity cycling, etc., must be addressed before the proposed novel
strengthening technique can be applied in practice. The authors noted that the method of
external reinforcement of concrete with advanced composites can be effective and economical,
not only in rehabilitation applications but also in new constructions.
Rashid et al. (2005) discussed the flexural test results of ten high-strength concrete beams
reinforced with aramid fiber-reinforced polymer (AFRP) bars. It was found that a concrete
beam, when reinforced with AFRP bars, became more flexible in the postcracking range than
an equivalent steel-reinforced beam, demonstrated wider and predominantly vertical cracks
even in the shear span, and may fail in an unusual flexure-shear mode. Cracks in AFRP-
reinforced beams formed in quick succession and, upon formation, penetrated deep into the
compression side of the beam straight away.
18
Arduini and Nanni (1997) conducted experimental testing and analytical analysis for the
case of beams precracked and subsequently strengthened with CFRP sheets. The authors
concluded that the strengthening technology consisting of externally bonded CFRP sheets was
easy to perform and resulted in improvements in ultimate load capacity and, to lesser extent,
in flexural stiffness. The area in need of major attention and, possibly, improvement was that
of concrete-FRP adhesion. It was necessary to avoid or at least limit the extent of FRP peeling
in order to improve the effectiveness of the strengthening method and the ductility of the load-
deflection response. It was possible to simulate and predict experimental load versus deflection
behavior, strain distribution, and the failure mode of FRP strengthened beams, including the
effects of pre-cracking and unloading-reloading cycles.
Bonacci & Maalej (2001) studied the performance of conventionally reinforced concrete
(RC) beams strengthened in flexure with externally bonded fiber-reinforced polymers (EB-
FRP) through compiling and analyzing an experimental database. A total of 127 specimens
from 23 separate studies were included in the database. The results revealed that one-third of
the specimens with added external reinforcement showed strength increases of 50% or more
in combination with considerable deflection capacity. It was concluded that future research on
the application of FRP to RC members should focus on conditions that were similar to what
was observed in the field, including the effects of sustained load during repair/strengthening
as well as corrosion- and load-induced damage. This is to assess the real potential of using FRP
for expedient and economical field repair and strengthening of RC members.
Ebead & Marzouk (2005) presented a tension-stiffening model for FRP-strengthened
concrete and a finite element analysis. It was found that a distinction of tensile behavior had to
be made between the plain, reinforced and strengthened reinforced concrete when defining the
tension-stiffening model. FRP-strengthened concrete exhibited a stiffer postpeak response
than conventional reinforced concrete. The use of CFRP strips and GFRP laminates were
sufficient to achieve positive results for flexural-strengthening of slabs. The strengthened
specimens using FRP strips or laminates showed an average gain in the load capacity of about
36% over that of the un-strengthened (control) specimens. In addition, the strengthened
specimens showed a stiffer behavior than that of the control specimens. However, a decrease
in ductility and energy absorption was recorded due to the brittle nature of the FRP composites.
For the suggested strengthening technique, de-bonding between FRP composites and concrete
19
was the major cause of failure. Slabs failed soon after de-bonding occurred due to exceeding
flexural capacity. None of the FRPstrengthened specimens experienced a rupture.
Mosallam & Mosalam (2003) presented an experimental and analytical investigation to
evaluate the ultimate response of unreinforced and reinforced concrete slabs repaired and
retrofitted with fiber reinforced polymer (FRP) composite strips. Both carbon/epoxy and E-
glass/epoxy composite systems were used in this study. It was found that the FRP systems
have succeeded in upgrading the structural capacity of both two-way unreinforced and
reinforced concrete slabs. For repair applications of unreinforced concrete slabs, test results
indicated that the composite system restored not only the original capacity of the damaged
slabs but also resulted in an significant increase of the strength of the repaired slabs to an
average increase of more than 540% the original capacity of the as-built slabs. For retrofitting
applications, the use of FRP systems resulted in noticeable upgrade of the structural capacity
of the as-built slabs up to 500% for unreinforced specimens and 200% for steel reinforced
specimens.
Maaddawy & Soudki (2008) examined the potential use of mechanically-anchored
unbonded fiber-reinforced polymer (MA-UFRP) system to upgrade the flexural strength of
deficient reinforced concrete (RC) slabs. The structural performance of slabs strengthened with
MA-UFRP system was evaluated and compared to that of slabs strengthened with externally-
bonded FRP (EB-FRP) system. All slabs had 0.8% steel reinforcement ratio while strengthened
slabs had an additional 0.12% external CFRP reinforcement ratio. Test results exhibited that
MA-UFRP system resulted in up to 43% improvement in the slab flexural strength. The
strength of the slabs strengthened with MA-UFRP system was 18% lower (on average) than
that of the slab strengthened with EB-FRP system with end-anchorage, but only 10% lower
than that of the slab strengthened with EB-FRP system without end-anchorage. The mid-span
deflection at ultimate load of the slabs strengthened with MA-UFRP system was 56% higher
(on average) than that of the slab strengthened with EB- FRP without end-anchorage, 5%
higher than that of the slab strengthened with EB-FRP with end-anchorage, and only 15%
lower than that of the control specimen.
Smith & Kim (2009) reported on the test results of one-way spanning reinforced concrete
(RC) slabs (with or without central cutouts) strengthened with fiber-reinforced polymer (FRP).
All FRP-strengthened slabs obtained a higher load-carrying capacity than their unstrengthened
20
(control) specimens. In addition, all strengthened slabs failed by debonding initiating at
intermediate cracks (a.k.a. IC debonding) and in the case of the slabs with central cutouts, the
critical cracks were diagonal and originated from the corners of the cutout. The extent of
debonding and the ability of the slab to sustain load post-initiation of debonding was dependent
on the position of the load. The slab in which the line load was located adjacent to the cutout
exhibited transverse bending action and as a result was able to withstand more extensive
debonding prior to loss of load-carrying enhancement from the FRP.
Teng et al. (2000) presented an investigation into the feasibility and effectiveness of
bonding GFRP strips to the top (tension) surface of deficient RC cantilever slabs as a
strengthening measure. The test results showed that a significant increase in the ultimate load
and ductility can be achieved if the slot anchorage system was used to anchor the strips into
the supporting wall. The effect of this strengthening method was even better if fiber anchors
were installed or the free ends of GFRP composite strips were wrapped around the free edge
and onto the soffit of the slab. They concluded that GFRP strips with slot anchorage to the
supporting wall and with fiber anchors to prevent debonding failure from the slab provided a
simple and effective system to strengthen deficient cantilever slab structures. The authors
recommended to address the following issues before this method can be widely used: (1) the
debonding mechanism, including its initiation and propagation; (2) the strength of the fiber
anchor in resisting combined tension and shear; (3) the capacity of the slot anchorage system;
(4) the effect of preloading; (5) a simple method to estimate the strength of slabs strengthened
using this method; and (6) durability of the strengthened slab.
Near Surface Mounted (NSM) FRP rods/bars
Khalifa (2016) studied the performance and failure mode of RC beams with NSM and EBR CFRP
strips, addressed the variables that may affect the flexural strength such as CFRP amount and
distribution and proposed a formula to compute the maximum strain for the NSM FRP material.
A set of beam specimens are tested and conclusions were drawn as follows:
· The ultimate load carrying capacity of beams strengthened with NSM CFRP strips was
higher than that for beams strengthened with EBR CFRP strips, due to the higher bond
strength of the CFRP strips in the case of NSM technique.
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· Increasing the amount of CFRP strips not necessarily results in a proportional increase in
the flexural capacity of the RC member especially if debonding of CFRP strips controls
the failure.
· Using the same amount of NSM CFRP Strips, and distributing the strips in two grooves
instead of one leads to a significant reduction in the crack widths and an increase in the
ultimate load.
· Failure of beams strengthened with EBR CFRP strips is controlled by debonding at the
strip–epoxy or the epoxy–concrete interfaces.
· For beams strengthened with NSM CFRP strips, failure is controlled by debonding of
CFRP strips and peeling off of the CFRP strips together with the concrete cover.
El-Gamal et al. (2016) studied the effect of technique used (NSM or Hybrid), type of FRP used
(carbon or glass), amount of FRP used, and steel reinforcement ratio on beam strengthening by
measuring the ultimate capacity, deflection, cracking, strains and mode of failure. A set of
specimens with different parameters were tested by four-point bending test. Based on the
experimental results, conclusions can be drawn as follows:
· All strengthened beams showed an increase in the ultimate capacity compared to the
reference beam. The strengthening was more efficient for the beams with lower steel
reinforcement ratios.
· NSM CFRP strengthened beams gave greater capacities than the NSM-GFRP strengthened
ones; however, they showed much more brittle behavior.
· The NSM-GFRP strengthened beam showed a very good ductile behavior with high
deflection values at ultimate load, which were almost similar or even greater than that
recorded in the reference beams. This gives ample warning before failure and can be
considered as an important advantage of using GFRP bars in the NSM strengthening
technique.
· Doubling the amount of the FRP strengthening material increased the ultimate capacity by
about 85%.
22
· The hybrid technique did not show an advantage compared to the NSM technique. In
addition, it needs more effort to prepare the surface of the beam before bonding the FRP
sheet. The NSM strengthening technique gave higher capacities (with two bars) and better
ductility (in case of using GFRP bars).
Sharaky et al. (2015) studied the effect of bond length of NSM reinforcement, construction
details and fiber reinforced polymer (FRP) characteristics on flexural performance. Based on the
experimental results, conclusions can be drawn as follows:
· Beams strengthened with fully bonded NSM FRP had a higher stiffness and bearing
capacity than those with partially bonded NSM FRP. The failure of the beams strengthened
with NSM CFRP and GFRP bars was concrete cover separation starting at the cutoff,
except in the beams with CFRP strips, which failed at the strip-epoxy interface.
· Increasing the bond length from 480 mm to 1000 mm increased the yield load. However,
there was no increase in the yield load when the bond length increased from 384 to 480 mm.
On the other hand, for beams where transverse wrapping and end anchorage were applied,
the maximum load can be increased.
· The deflection of the beams strengthened with fully bonded NSM bars was lower than for
those with partially bonded NSM. Also, there is only a slight effect of bond length on
deflection in the partially bonded NSM beams. For the beams strengthened with NSM
strips, the same behavior was observed.
· The strain gauges in the constant moment region should have had almost the same readings.
However, the results for beams were unexpected due to instability in the electrical strain
gauge readings during the failure stages.
Bilotta et al. (2015) studied the flexural behavior of beams strengthened with NSM and EBR
techniques. Moreover, it investigated the effect of the loading pattern on failure modes due to
different distributions of bending moment and shear along the beam under different loading
patterns (concentrated load and distributed load). Based on experimental results, conclusion can
be drawn as follows:
23
· Despite the lower transversal area, the maximum loads of both beams strengthened with 2
or 3 NSM strips is comparable with the ones achieved by the EBR beams under both
loading schemes. Therefore, the efficiency of NSM is confirmed to be higher than EBR
systems.
· The NSM strips are less sensitive to debonding phenomena, because (a) debonding at the
intermediate crack did not occur, (b) a shear failure of the beam occurred when the shear
at the anchorage zone becomes very high or (c) the formation of a Critical Diagonal Crack
(CDC), starting from the application point of the concentrated loads, led to debonding of
strips with Concrete Cover Separation (CCS).
· The NSM strips are less effective in increasing the RC beam stiffness than EBR plates,
when only two strips are applied.
· For EBR beams, the End Debonding failure mode is more critical when they are tested
under distributed loads because the shear attainable at the anchorage is higher.
· The beams with NSM tested under concentrated loads failed due to Critical Diagonal Crack
Debonding with Concrete Cover Separation, while in the beams tested under distributed
loads, a shear failure occurred instead of debonding. Under distributed load, indeed, the
shear at the ends of the beam is greater than in the case of the concentrated loads and,
therefore, the critical failure mode can move from bending to shear if debonding
phenomena do not occur.
Wang et al. (2015) conducted a comprehensive study on RC beams prestressed by basalt FRP
(BFRP). He mainly studied the effects of tension stress and the tendon profile of the prestressing
tendons on the flexural behavior of RC beams, including cracking load, yielding load, ultimate
load, stiffness and ductility. The determination of tensile stress is based on the creep rupture
limitation (0.52 fu is adopted in this paper). According to the experiment and FE simulation,
conclusions are drawn as follows:
· the tensile stress limit of 0.5fu, a deviator angle of 2 degree and bonding anchors satisfy
requirements for safety and efficiency.
· BFRP tendons can increase the cracking load, yielding load, ultimate load regardless of
tension stress and tendon profile. Compared to straight tendon profile, deviated BFRP
24
tendons can benefit more from the increase in the cracking, yielding and ultimate load.
· a higher tension stress of BFRP tendons can result in higher structural performance.
Notes: Because of high creep rupture stress in comparison with GFRP(tensile stress is less than
0.3fu), CFRP and AFRP are adopted as prestressing mterials. But the high cost and brittle
behavior of CFRP and large relaxation and cost issue for AFRP limit their application.
Sharaky et al. (2014) studied the behavior of RC beams strengthened with NSM FRP bars of a
limited bond length. The study mainly focused on the effect of FRP material (carbon or glass),
the number of NSM bars and their area, epoxy properties, and the strengthening arrangement on
the flexural behavior. The load capacity, deflection, mode of failure, FRP strain, concrete strain,
free end slip and transverse strain in epoxy and concrete were analyzed. Based on the four-point
bending test, conclusions can be drawn as follows:
· Compared to beams strengthened with GFRP bars, CFRP can increase the larger yielding
loads, depending mainly on the area of the FRP bars and the epoxy properties.
· Increasing the number of NSM CFRP and GFRP bars increased the yielding and the
maximum loads. However, the deflection, crack width and length of beam showed the
opposite trend. Moreover, the epoxy type had no effect on the strengthened beams’
stiffness.
· The recorded end slips for the two beams each with two NSM bars were slightly higher
than those of the beam with one NSM bar, due to a lower confinement (edge effect) in the
case of two NSM bars, while the type of FRP bars had little effect on the end slip.
· The NSM technique is effective for increasing the load capacity and stiffness of RC beams.
The technique’s load efficiency depends mainly on the area of the FRP bars and their mode
of failure, while stiffness enhancement is mainly influenced by the bars’ modulus of
elasticity.
Rezazadeh et al. (2014) studied the influence of the prestressing technique on the flexural
behavior of RC beams strengthened with NSM CFRP laminates. Based on the experimental
results, conclusions can be drawn as follows:
25
· After releasing the prestress force, a negative camber is generated due to the negative
bending moment caused by the eccentricity of this force in relation to the centroidal axis
of the beam’s cross section. This negative camber resulted in a decrease of tensile strain in
the prestressed CFRP laminate, which represents the short-term prestress loss immediately
after release.
· The prestress force created an initial compressive strain in the tensile steel reinforcement
and surrounding concrete, which led to an increase of the load carrying capacity at concrete
cracking and steel yielding initiations.
· Based on the results obtained for the control beam, all CFRP strengthened beams provided
an increase in terms of ultimate load carrying capacity, regardless the prestress level
applied to the CFRP laminate.
· The load carrying capacity at serviceability limit conditions increased significantly with
the prestress level, when compared to the load carrying capacity of the beam strengthened
with a non-prestressed CFRP laminate. However, the ultimate deflection of the
strengthened beams decreased with the increase of applied prestress level. These results,
which imply a decrease of the ductility index with the prestress level, suggest the adoption
of an upper limit of the prestress level to be applied to the CFRP laminates in order do not
compromise the ductility performance of the RC beams strengthened according to the
proposed technique.
· All strengthened beams failed by rupture of the CFRP laminate after the yielding of the
tension steel reinforcement. The results showed that the possibility of the concrete crushing
as the prevailing failure mode decreases with the increase of the prestress level applied to
the CFRP laminate.
· The crack patterns of all beams consisted predominantly of flexural cracks. The cracked
zone length of the beam strengthened with a non-prestressed NSM CFRP laminate
increased when compared to the control beam, while the increase of the prestress level
inverted this tendency, by decreasing the cracked zone length in comparison to the passive
strengthened beam. The strengthening system, regardless the prestress level, has also
provided a decrease of average crack spacing when compared to the control beam.
26
Almassri et al. (2014) studied the validity of a repair technique using NSM CFRP rods to restore
the mechanical performance of corrosion-damaged RC beams. Based on the three-point testing,
conclusions can be drawn as follows:
· The NSM technique can increase the ultimate load capacity of a corroded beam that has
suffered considerable damage and can allow it to reach to the ultimate capacity of the
control beam.
· The efficiency of the NSM technique in repairing corroded beams could be limited by the
separation of the concrete cover due to corrosion induced cracks.
· The NSM technique slightly increases the stiffness of both repaired corroded and repaired
control beams.
· The NSM technique increases the ultimate deflection value for repaired control and
corroded beams.
· The NSM technique restores sufficient ductility after ductility loss due to the brittle
behavior of corroded RC beams because of steel corrosion.
· If there is 1% cross-section loss (it is computed by vernier caliper or weight loss) due to
steel corrosion it will be reflected as a 1% loss in the yielding capacity value. The
percentage is different for ultimate capacity as the mode of failure is not the same in each
case.
De Lorenzis et al. (2002) investigated the mechanics of bond between NSM FRP rods and
concrete through a newly developed specimen with the advantages of the direct pull-out type
of test while minimizing the problem of eccentricity, the preparation time and the use of
material. They analyzed the influence of the most critical parameters on the bond performance
including type of FRP rod (material and surface pattern), groove-filling material, bonded
length, and groove size. For specimens with epoxy resin and spirally wound or ribbed CFRP
rods, failure at the epoxy–concrete interface was the critical mechanism in all cases, due to the
smooth surface of the grooves. Specimens with ribbed GFRP rods experienced a shift in failure
mode as the depth of the groove increased, passing from splitting of the epoxy cover,
27
accompanied by cracking of the concrete surrounding the groove, to failure at the epoxy–
concrete interface. For specimens with cement mortar, splitting of the cover was more frequent
than for epoxy, due to the lower tensile strength of the material. However, the ultimate load of
these specimens was in all cases lower than that of epoxy-filled specimens.
Khalifa et al. (1999) developed an innovative anchor system called U-anchor to allow a
better exploitation of the potential of strengthening and rehabilitation technique for RC
structural elements with externally bonded FRP composites. The U-anchor system provideed
an effective solution for cases in which the bonded length of FRP composites was not sufficient
to develop its full capacity or where anchorage to adjacent members was required. The U-
anchor can be used with FRP sheets and pre-cured laminates that were unbonded or fully
bonded to concrete. An example of experimental verification was discussed to show the
feasibility and effectiveness of the proposed system in increasing the shear capacity of RC
beams strengthened with CFRP U-wraps. For a beam strengthened with CFRP without U-
anchor, shear capacity increased but failure was governed by debonding of the CFRP. In the
specimen where the anchor was used, shear capacity of the member further increased and no
FRP debonding was observed at ultimate. The authors concluded that there was sufficient
evidence to indicate that the novel U-anchor will make FRP strengthening even more attractive
and economical for the concrete/masonry repair industry.
Barros et al. (2005) carried out series of tests with concrete columns, concrete beams and
masonry panels to evaluate the effectiveness of a near surface mounted (NSM) strengthening
technique for elements failing in bending and elements failing in shear. The NSM technique
was based on bonding laminate strips of CFRP into slits made onto the concrete cover of the
elements to be strengthened. The results showed that the proposed strengthening technique was
very promising for increasing the load carrying capacity of concrete columns failing in
bending. The NSM technique was also very effective to increase the flexural resistance of RC
beams. It provided a higher increase on the load-carrying capacity of masonry panels as well
as a larger deflection at the failure of the panels. The authors found that the NSM technique
was easier and faster to apply than externally bonded reinforcing (EBR) technique.
Nordin & Täljsten (2006) presented an experimental investigation on prestressed CFRP
quadratic rods bonded in sawed grooves in the concrete cover. This method has proven to be
an advantageous means of bonding CFRP to concrete. The shear and normal stress between
28
the CFRP and the concrete were more efficiently transferred to the structure in comparison to
surface bonded laminates. In their experiment, no mechanical device has been used to maintain
the prestress during testing, which means that the adhesive must transfer all shear stresses to
the concrete. Fifteen beams with a length of 4 m were experimentally tested. Test results
showed that the prestressed beams exhibited a higher first-crack load as well as a higher steel-
yielding load as compared to nonprestressed strengthened beams. The ultimate load at failure
was also higher, as compared to nonprestressed beams. In addition, the beams strengthened
with prestressed FRP had a smaller midpoint deflection. All strengthened beams failed due to
fiber rupture of the FRP.
Turco et al. (2006) presented experimental results of different applications of NSM FRP
bars for the strengthening of masonry walls. Each of them showed promising potential for
retrofitting of masonry structures. In the case of flexural strengthening, the capacity was
increased by a factor of up to 2.5 and by 4.5–26 times in the case of shear strengthening. The
glass FRP has proved to be a good material for masonry strengthening. In spite of its low elastic
modulus, it usually performs better than the carbon FRP. The authors found that smooth
circular FRP bars with low bonding capacity were appropriate for shear strengthening, while
rectangular FRP bars exhibited good performances in the case of flexural strengthening. Low-
bond systems (i.e. sand-coated FRP bars + cementitious paste, smooth FRP bars + epoxy paste)
were preferable in the case of shear strengthening as they allowed some sliding and a better
redistribution of the stresses in the system. Similar results were achieved by using epoxy and
cementitious paste as embedding material. However, since cementitious paste was cheaper and
preserved better the appearance of the original wall, it was more attractive and promising for
retrofitting of existing masonry structures.
Bournas and Triantafillou (2009) presented the results of a large-scale experimental
program aiming to study the behavior of reinforced concrete (RC) columns under simulated
seismic loading, strengthened in flexure with different types and configurations of near-
surface-mounted (NSM) reinforcing materials. Test parameters included carbon or glass fiber-
reinforced polymers (FRP) versus stainless steel, configuration and amount of NSM
reinforcement, confinement via local jacketing, and type of bonding agent (epoxy resin or
mortar). The results demonstrated that NSM FRP and stainless steel reinforcement was a viable
solution toward enhancing the flexural resistance of RC columns subjected to seismic loads.
29
This was especially the case when the retrofitting scheme combined epoxy-bonded NSM bars
with local confining jackets (textile-reinforced mortars as in this study).
Hassan & Rizkalla (2003) presented experimental and analytical investigations to evaluate
bond characteristics of near surface mounted carbon FRP (CFRP) strips. The results showed
that the use of near surface mounted CFRP strips was feasible and effective for
strengthening/repair of concrete structures. The use of near surface mounted CFRP strips
substantially increased both stiffness and strength of concrete beams. The authors found that
debonding loads increased by increasing the embedment length of CFRP strips, concrete
compressive strength, and/or groove width. In addition, development length of near surface
mounted CFRP strips increased by increasing the internal steel reinforcement ratio. The
development length decreased with the increase of either the concrete compressive strength
and/or the groove width.
Barros et al. (2016) carried out an experimental program composed of a series of RC beams
of relatively high T cross section (600 mm height) that was shear-strengthened with CFRP
laminates using NSM technique. The laminates were positioned at different depths into slits
opened on the side faces of the beam’s web. They concluded that the NSM technique with
CFRP laminates was very effective in RC beams of relatively high cross section, not only in
terms of increasing the overall behavior of the RC beams, but also in assuring higher
mobilization of the tensile properties of the CFRP. The CFRP shear strengthening
configurations provided an increase in terms of maximum load that ranged between 66% and
81%. The authors found that the deeper the laminates were installed into the beam’s web, the
higher shear strengthening effectiveness they can provide. The shear strengthening
configuration constituted by two independent laminates in the same slit, as well as the
strengthening configuration where laminates of two distinct inclinations (52 and 90 degrees)
were installed at different depth inside the slit, have provided a more ductile behavior after
peak load for the beams shear strengthened with these configurations.
Barros & Dias (2006) conducted an experimental program of four-point bending tests to
assess the effectiveness of near surface mounted (NSM) technique for the shear strengthening
of concrete beams. Four beam series of distinct depth and longitudinal tensile steel
reinforcement ratio were tested. Each series was composed of one beam without any shear
reinforcement and one beam using the following shear reinforcing systems: (1) conventional
30
steel stirrups; (2) strips of wet lay-up CFRP sheet embracing the bottom (tension side) and the
lateral beam faces, designated by externally bonded reinforcing (EBR) technique; and (3)
laminates of CFRP embedded into vertical or 45° inclined pre-cut slits on the concrete cover
of the beam lateral faces (NSM strengthening technique).
2.3.2. Shear Strengthening
Baggio et al. (2014) studied the effectiveness of using CFRP, GFRP and fiber reinforced
cementitious matrix (FRCM) sheets to increase the shear capacity of RC shear critical beams and
investigated the effect of presence and type of FRP anchors (CFRP or GFRP) on shear capacity.
Conclusions drawn from the test results were as follows:
· All beams exhibited a similar flexural stiffness during testing regardless of the
strengthening material, properties and layout (depth or width).
· Beams strengthened with FRCM showed a 31% and 34% increase in shear capacity with
and without anchors over the control specimen, respectively, and did not result in a change
in the mode of failure.
· Strengthening using full depth GFRP sheets and no anchors resulted in a 50% increase in
shear capacity over the control with debonding occurring before diagonal tension shear
failure.
· Beams strengthened with partial depth GFRP sheets with and without anchors showed a
52% and 36% increase in shear capacity, respectively, over the control beam and the
presence of the FRP anchors was effective in halting the debonding of the GFRP sheets.
· Beams strengthened with CFRP sheets with and without anchors showed a 67% and 75%
increase in shear capacity over the control beam, respectively, with both beams exhibiting
a ductile flexural failure mode.
· CFRP strengthening changed the mode of failure from a brittle shear failure to a flexure
failure by yielding of the longitudinal steel rebar.
· The use of GFRP anchors provided an additional 13% increase in strength over the control
compared to the equivalent GFRP-strengthened beam with CFRP anchors.
31
· Although the presence of FRP anchors halted the debonding process, the anchored FRP
sheets did not reach their ultimate rupture strain.
· When the available bonded length is limited, the installation of FRP anchors is a viable
option to prevent a brittle shear failure mode due to FRP debonding.
El-Sayed (2014) studied the shear strength of beams strengthened for flexure with EBR-CFRP
reinforcement and investigated the effect of steel reinforcement ratio and with/without stirrups on
the shear capacity. Main findings of this investigation can be summarized as follows:
· All CFRP-strengthened beams showed higher shear strength in comparison to that of the
beams without external reinforcement. The increase in shear strength was up to 35%. This
finding reveals that external CFRP longitudinal strengthening contributes to the concrete
shear strength of reinforced concrete beams.
· The effect of FRP longitudinal strengthening on the shear strength can be captured by
combining the reinforcement ratios of internal steel and FRP in an equivalent
reinforcement ratio.
Koutas and Triantafillou (2012) presented an experimental investigation on the
effectiveness of various types of spike anchors in combination with U-shaped fiber-reinforced
polymer (FRP) jackets for shear strengthening of reinforced concrete T-beams. The parameters
examined included the orientation, the number and spacing of anchors, and the role of carbon
versus glass fibers in the anchors. They concluded that anchors placed inside the slab were
many times more effective than those placed horizontally inside the web, and anchors of
similar geometrical characteristics (e.g., embedment length) exhibited similar effectiveness
despite the difference in fiber type.
Monti (2007) developed a mechanics-based (as opposed to regression-based) model of the
shear capacity of reinforced concrete beams, strengthened with externally bonded fiber-
reinforced polymers (FRP). The model was obtained through three major steps including: (1)
First, the generalized constitutive law of an FRP layer bonded to concrete was defined; (2)
Second, the compatibility imposed by the shear crack opening and the appropriate boundary
conditions – which depended on the strengthening configuration (either side bonding, U-
jacketing or wrapping) – were included in the formulation; and (3) Finally, analytical
expressions of the stress field in the FRP strip/sheet crossing a shear crack were obtained.
32
Through these expressions, closed-form equations for the effective debonding strength of FRP
strips/sheets were defined as function of, both, the adopted strengthening configuration, and of
some basic geometric and mechanical parameters. The developed equations exhibited good
correlation with his experimental testing and with test data collected from the literature.
Micelli et al. (2002) presented the results of an experimental investigation on twelve
reinforced concrete (RC) T-joists strengthened with fiber-reinforced plastic (FRP) composites.
Different strengthening schemes, including different FRP materials and a new FRP anchorage
system, were evaluated. Carbon FRP and aramid FRP sheets in an epoxy matrix were bonded
to the RC joists using the wet layup technique. All of the unanchored FRP strengthened beams
showed failure due to peeling, while the anchored FRP strengthened members showed failure
due to anchor pullout at higher load values. It was found that an increase in the amount of FRP
did not result in a proportional increase in the shear capacity, as expected by design equations,
but all of the beams showed a considerable increase in stiffness.
Adhikary and Mutsuyoshi (2004) presented an experimental testing program on shear
strengthening of continuous unidirectional flexible carbon-fiber tow sheets (hereafter CFS)
bonded to reinforced concrete (RC) beams. It was found that the externally adhesive bonded
flexible CFS can increase the ultimate shear strength of RC beams and enhance the flexural
stiffness. This method can be used effectively for shear strengthening or upgrading RC beams
as it was relatively easy for construction and handling. Maximum shear strength was obtained
for the beam with full U-wrapped sheets having vertically aligned fibers. Beams bonded with
FRP sheets having horizontally aligned fibers also showed enhanced shear strengths as
compared to the control beam. CFS bonded beams showed substantially delayed diagonal
cracking. This result affirmed the sheet’s contribution to shear strength prior to the occurrence
of a diagonal crack.
2.3.3. Column Strengthening and Confinements
The noticeable benefits of upgrading concrete columns with FRP composites included
increased ductility (as a result of the confinement provided by FRP wraps, the concrete will
fail at a larger strain than unconfined columns), increased strength (the lateral pressure exerted
by the wraps will increase the compressive strength of the concrete in both the core and shell
regions, resulting in higher load-carrying capacity), flexibility of the FRP wraps (allows
33
wrapping around circular as well as rectangular and square columns), low weight, low
maintenance, aesthetic appearance (FRP wraps are very thin), and can be used both as
permanent or temporary solution (if more effective alternatives are developed down the road,
the FRP wraps can be easily removed) (Saadatmanesh et al. 1994).
Parghi and Alam (2016) studied the effects of Material properties, amount of longitudinal and
transverse steel, external confinement, axial load and shear span-depth ratio on the limit states of
carbon fiber reinforced polymer (CFRP) confined seismically deficient RC circular bridge piers
using fractional factorial design method. According to the experiment and FE modeling of piers,
conclusion can be drawn as follows:
· Height-to-depth ratio of column is the most important factor which affects the seismic
performance and limit states criteria of CFRP confined piers.
· Although the tie spacing does not show significant effect on the lateral load resistance
of a CFRP retrofitted pier, it shows some effect on the drift at yield, crushing and bar
buckling.
· The amount of longitudinal steel reinforcement shows significant effect on the yield
and crushing drift limit states and base shear of the CFRP confined bridge piers.
· Yield strength of longitudinal steel and interactions among longitudinal steel,
compressive strength, axial load and the shear span depth ratio significantly affect the
drift and base shear at yielding and crushing of the CFRP confined bridge piers.
· Compressive strength of concrete does not affect the yielding and crushing drift
significantly; however, it shows some contribution to the yielding and crushing base
shear to the seismic performance of CFRP confined bridge piers.
· Shear span-depth ratio and axial load level, and the interactions between yield strength
and concrete compressive strength, axial load and number of CFRP layers affect the
buckling and fracture of longitudinal steel reinforcements of the CFRP confined
deficient bridge piers.
Nisticò et al. (2014) assessed the performances of significant predictive expressions
concerning peak strength and ultimate strain based on more than 655 model tests published in
literature. There are mainly 2 ways to predict strength expressions, the classical approach and
linear function or nonlinear function of corner radius ratio (corner radius over the inscribed
34
circular section radius). The classical approach is the modification of the equation valid for
circular sections, through a coefficient that considers the reduction of the confinement efficiency.
It divides the square section in a zone confined by a confinement pressure equal to that of the
corresponding circular section and in an unconfined zone. The coefficient, therefore, is equal to
the ratio of effectively confined area divided by the total section area. Two parameters are used
to assess existing models: the Average Absolute Error (AAE) and the Average Ratio (AR).
According to two parameters, we can conclude that errors related to peak strength are lower than
that of ultimate strain. The more suitable expression for the assessment of fcc is
Rousakis
et al.
The suitable ultimate strain expression for circular sections is
De Lorenzis and Tepfers (2003)
Rousakis et al. (2012)
The suitable ultimate strain expression for circular and square sections is
CNR DT 200/2004 (2004)
Wei and Wu (2012)
Parvin et al. (2014) draw some conclusions based on the review of FRP-strengthening of
columns for several loading scenarios including impact load.
35
· The angle and hoop plies and stacking sequence in wrap configuration provided different
level of ductility and strength for the columns with identical FRP wrap thickness.
· Application of hoop and ply combination for wrap configurations on prismatic columns
should be pursued, since they may delay premature fracture at the corners.
· Modifying the shape of square-to-circular and rectangular-to-elliptical columns will
eliminate the corner stress concentration in prisms and improve confinement effectiveness.
Subsequent FRP-wrapping of shape-modified columns will substaintially improve axial
load and pseudo ductility.
· More accurate and reliable models of confined concrete should be investigated through
comprehensive set of data for all column shapes to not only predict the strength but axial
and lateral strains as well.
· Lower strength concrete columns benefit the most in terms of compression load capacity
increase once confined with FRP sheets.
· For eccentrically loaded columns, smaller enhancement factor should be considered in
design of FRP-wrapped concrete columns.
· Seiemic damage to deficient RC columns can be reduced or completely prevented by
applying unidirectional fiber composite sheet along the longitudinal direction to increase
flexural capacity, and by wrapping the columns in the lateral direction to improve their
ductility and energy absorption capacity.
· To withstand impact loadings, concrete columns should be properly strengthened to
achieve adequate level of energy absorption capacity and ductility.
· The FRP repair of corrosion damaged RC columns not only provides strength and ductility,
but also could slow down the rate of the corrosion reaction.
Siddiqui et al. (2014) studied the effectiveness of hoop and longitudinal Carbon FRP (CFRP)
wraps in reducing the lateral deflections and improving the strength of slender circular RC
columns (slender columns take second-order into account). According to the experiment,
conclusions were drawn as follows:
· CFRP hoop wraps provide confinement to concrete and lateral support to the longitudinal
fibers and thus increase the strength of both short and slender RC columns. However, the
effect of hoop wraps on the strength of columns is more significant for short columns than
36
slender columns. The outer hoop layer in slender columns also helps in reducing the
possibilities of flexural debonding of inner longitudinal FRP fibers from the concrete
surface.
· The fibers of longitudinal CFRP laminate can also contribute substantially to the load
carrying capacity provided they are laterally supported by hoop fibers. In slender columns,
the load is primarily carried by flexural action of longitudinal FRP fibers.
· The column strengthened with single layer of hoop FRP not only increases the load
carrying capacity but also the ductility substantially. For those specimens which were
strengthened using single layer of hoop FRP, the ductility increases with the length of
column. The increase in load capacity had the opposite trend.
· For those specimens which were strengthened with 2 or 4 layers of longitudinal and then
one layer of hoop FRP, the ductility and the ultimate load decrease with the length of
column.
· The increase in load carrying capacity through increase in the number of longitudinal CFRP
layers from two to four is substantially less compared to when it was increased from zero
to two.
Youssef et al. (2007) developed a stress–strain model for concrete confined by fiber
reinforced polymer (FRP) composites. The model was based on the results of a comprehensive
experimental program including large-scale circular, square and rectangular short columns
confined by carbon/epoxy and E-glass/epoxy jackets providing a wide range of confinement
ratios. It was found that the stress–strain curve for concrete confined by FRP composites
behaved bi-linearly. The first portion of the stress–strain curve traces that of unconfined
concrete until the jacket start getting activated. At this point, the curve became either ascend
or descend, depending on the geometry of the cross-section and the provided confinement ratio.
The proposed model was successful in covering a wide range of confinement ratios, which was
possible to generate by testing large-scale columns as opposed to the majority of the published
tests that were conducted on standard cylinders (152 mm × 305 mm or 6 in. × 12 in.).
Rousakis & Karabinis (2012) discussed experimental results of reinforced concrete square
section specimens, externally confined by carbon or glass FRP sheets. It was found that
37
external strengthening by composite materials through confinement can significantly enhance
both strength and strain ductility of such columns subjected to monotonic or cyclic loading.
Plain concrete columns strengthened by a jacket with Modified Confinement Ratio (MCR) =
0.184, revealed a slightly adequate mechanical behavior with softening branch. Columns with
adequate confinement leading to a hardening behavior displayed a higher improvement of the
strain at failure. In columns with slender bars, a softening branch resulted in inadequate load
response for MCR = 0.153 or lower. In those cases, the jacket cannot restrict the unstable
expansion of concrete core and to provide a hardening inelastic behavior. Both columns with
slender and non-slender bars had no significant variation of the axial strain at failure.
Pellegrino & Modena (2010) explored the behavior of reinforced concrete columns
confined with FRP sheets. In particular, some new insights on interaction mechanisms between
internal steel reinforcement and external FRP strengthening and their influence on efficiency
of FRP confinement technique were present. A procedure to generate the complete stress-strain
response was developed. New analytical proposals included: (1) effective confinement
pressure at failure; (2) peak stress; (3) ultimate stress; (4) ultimate axial strain; and (5) axial
strain corresponding to peak stress for FRP confined concrete elements with circular and
rectangular cross section, with and without internal steel reinforcement. Based on wide
database regarding FRP confined concrete columns, with circular and rectangular cross
section, with and without existing steel reinforcement, the authors concluded that the new
proposal appeared to be the most accurate, compared to a number of analytical models
available in literature.
2.4. Upgrading metallic structures
Jones & Civjan (2003) investigated the effectiveness of applying CFRP overlays to steel
fatigue tension coupons to prolong their fatigue life. Specimens were either notched or center-
hole specimens and tested in uniaxial tension. Test variables included CFRP system, bond
length, bond area, one and two sided applications, and applications prior or subsequent to crack
propagation. Two sided applications were very effective, prolonging fatigue life by as much
as 115%. Similar application of CFRP materials subsequent to crack propagation extended the
remaining fatigue life by approximately 170% without any other means of crack arrest. It was
found that the use of the CFRP overlays is promising as both a preventive technique and repair
38
method. The epoxy performance was critical to the effectiveness of the system, with all failures
initiated by debonding of the CFRP. Overlays were most effective when the system was
applied directly to the potential crack trajectory. One-sided applications introduced bending
and did not significantly improve performance. CFRP materials with a moderate modulus of
elasticity performed best.
Shaat & Fam (2006) presented the behaviour of axially loaded short and long square hollow
structural section (HSS) steel columns, strengthened with CFRP sheets. Twenty-seven short-
column and five long-column HSS specimens were tested. The effect of CFRP sheet
orientation in the longitudinal and transverse directions was studied for short columns. For
long columns, CFRP sheets were oriented in the longitudinal direction only. A maximum
strength gain of 18% was obtained for short columns with two transverse CFRP layers. For
long columns, the maximum strength gain of 23% was achieved with three longitudinal CFRP
layers applied on four sides. In all CFRP-strengthened long columns, lateral deflections were
reduced. Strength gain in long columns was highly dependent on the column’s imperfection.
Teng & Hu (2007) explored the use of FRP confinement to enhance the ductility and hence
the seismic resistance of circular steel tubes. A series of axial compression tests has been
conducted to evaluate the effectiveness of FRP confinement of steel tubes whose ductility was
limited by the development of the elephant’s foot buckling mode. A finite element model for
predicting the behaviour of these FRP-confined tubes has also been presented. Both
experimental and numerical results have shown that with the provision of a thin FRP jacket,
the ductility of the steel tube can be greatly enhanced. These results have also exhibited that
when the jacket thickness reached a threshold value for which inward buckling deformations
dominated the behaviour; further increases in the jacket thickness did not lead to significant
additional benefits as the jacket provided little resistance to inward buckling deformations. It
is important to note that FRP confinement of steel tubes leads to large increases in ductility but
limited increases in the ultimate load, which is desirable in seismic retrofit so that the retrofitted
tube will not attract forces, which are so high that adjacent members may be put in danger.
2.5. Upgrading timber and masonry structures
Borri et al. (2005) described an experimental and numerical study on the reinforcement of
existing wood elements under bending loads through the use of FRP materials. Mechanical
39
tests on the reinforced wood exhibited that external bonding of FRP materials may produce
increases in flexural stiffness and capacity. The FRP composite material was made of high
tensile carbon mono-directional reinforcing fabrics embedded in an epoxy resin matrix. This
reinforcing method can be applied without necessitating the removal of the overhanging part
of the pre-existing wood structure, thus maintaining the original historical structure. The
developed numerical model based on Bazan-Buchanan law has shown that it was possible to
estimate the failure loads of a FRP-reinforced wood element.
Li et al. (2014) proposed hollow-sectioned wood beams strengthened by GFRP rod and
CFRP composite sheets. Five groups of specimens with and without FRP reinforcement were
tested. The four-point bending test was conducted to obtain the force–displacement
relationships of the wood beams. Test results showed that the average strengths of the
strengthened wood beams were increased approximately 4.3%-9.5% compared to
unstrengthened (control) beams.
Raftery & Kelly (2015) discussed the results of an experimental program to examine the
use of basalt fiber-reinforced polymer rods for the strengthen air of low-grade glued laminated
timber in flexure. It was concluded that the basalt FRP rods can effectively bridge over
damaged zones in the timber and comprehensively restore the mechanical strength and
stiffness of the original undamaged section. Unreinforced beams generally showed brittle
tensile failures with fractures initiating at defects in the bottom laminations. Reinforced and
artificially fractured repaired beams which used bonded-in basalt FRP rods in the tension zone
generally showed considerable ductility with visible compression wrinkling in the top
laminations. The amount of ductility that was experienced by the reinforced sections was
influenced by the distance of the reinforcing rods from the neutral axis. The authors suggested
that with increasing acceptance of basalt FRP profiles in the construction industry, there existed
considerable potential for the development of sustainable basalt-timber hybrid elements.
Altin et al. (2008) presented an experimental investigation on the behavior of 1/3 scale,
one-bay, one-storey nonductile RC frames with masonry infill walls strengthened by diagonal
CFRP strips and subjected to cyclic lateral loading. The diagonal CFRP strips that were used
to retrofit brick masonry infilled reinforced concrete frames were effective to significantly
increase the lateral strength and stiffness of the frames. An important advantage of this
technique was that it can be applied without evacuating the building during its application, thus
40
causing minimum disturbance to the occupants. Lateral strengths of specimens strengthened
with the CFRP strips on both sides of the infill wall increased by 2.18 and 2.61 times compared
to specimens receiving CFRP strips on one side only. The authors found that when the width
of CFRP strip was increased, increase in strength and stiffness could be limited.
Li et al. (2005) developed an analytical model to predict the effectiveness of fiber-
reinforced polymer (FRP) composite materials in retrofitting unreinforced masonry (URM)
walls to reduce seismic damage. The reinforcement considered was near-surface-mounted
(NSM) FRP bars. Test results exhibited that the FRP strengthening technique can be very
effective in significantly increasing the in-plane strength and ductility of URM walls. The
maximum increment in shear capacity was approximately 80%, registered in walls
strengthened with GFRP bars placed at every bed joint. Compared with unreinforced walls,
strengthened walls showed much better pseudo-ductility and were much more stable after
failure (that is, no loss of material was observed), potentially reducing the risk of partial or
total collapse. Walls strengthened with the same amount of reinforcement showed similar
performance.
2.6. NDT/NDE process for inspecting FRP wraps
Taillade et al. (2011) presented the evaluation of the adhesive bond between concrete and
externally bonded FRP on concrete surface structures using two techniques including
shearography and pulsed stimulated infrared thermography. Results showed that shearography
associated to the depressure loading made it possible to determine not only locations and areas
of defects but also enabled to assess quality of adhesion. Moreover, thermography offered a
simple method with real-time and full-field imaging capabilities that permited to inspect
repaired structures in a qualitative way (detection of the bonding defects).
Saafi & Sayyah (2001) developed an active damage interrogation system which used an
array of piezoelectric transducers attached to or embedded within the structure to monitor the
health of concrete structures repaired and reinforced with composite materials. The results
demonstrated the ability of active damage interrogation to detect, identify, and localize
disbands and delamination in concrete externally reinforced with carbon FRP. Delaminations
in concrete beam reinforced with composite materials subjected to loading were successfully
identified. The damage indices obtained from the piezoelectric transducers were confirmed by
41
the experimental observations in terms of crack propagation and the effects on the
delamination. It was found that in each zone, the damage index increased with increasing
applied load. The authors concluded that the active damage interrogation method appeared to
be a promising method for assessing the damage and its severity in concrete repaired with
composite materials.
2.7. National guidelines pertaining to FRP-retrofit
(a) Design guidelines based on the ACI 440.2R-08
1. Materials and Properties
Materials
· Resins- Epoxy, vinyl esters, and polyesters are commonly used resin types.
· Adhesives- Adhesives are used to bond procured FRP laminate and NSM systems
to the concrete substrate.
· Fibers- Glass, aramid, and carbon fibers are common reinforcements used with FRP
systems.
Physical properties
· Density- Reduced density leads to lower transportation costs, reduces dead load on
the structure, and can ease handling of the materials on the project site.
· Coefficient of thermal expansion- The coefficients of thermal expansion of
unidirectional FRP materials differ in the longitudinal and transverse directions。
· Effects of high temperatures- Beyond the Tg, the elastic modulus of a polymer is
significantly reduced due to changes in its molecular structure.
Mechanical properties
· Tensile behavior- The tensile behavior of FRP materials is characterized by a linear
elastic stress-strain relationship until failure, which is sudden and brittle.
· Compressive behavior—Externally bonded FRP systems should not be used as
compression reinforcement due to insufficient testing validating its use.
Table 1- Typical densities of FRP materials, lb/ft^3
42
Steel GFRP CFRP AFRP
490 75 to 130 90 to 100 75 to 90
2. Design Considerations
Design limit states
· Strengthening limit states- The unstrengthened structural member
without FRP reinforcement, should have sufficient strength to resist a
certain level of load.
(𝜑𝑅𝑛) 𝑒𝑥𝑖𝑠𝑡𝑖𝑛𝑔 ≥ (1.1𝑆𝐷𝐿 + 0.75𝑆𝐿𝐿) 𝑛𝑒𝑤
· Structural fire endurance- The nominal strength at high temperature
should be greater than the strengthened service load on the member.
𝑅𝑛 ≥ 𝑆𝐷𝐿 + 𝑆𝐿𝐿
Design material properties
· Material properties used in design equations should be reduced based
on the environmental exposure condition.
f𝑓𝑢 = C𝐸 f′𝑓𝑢, Ɛ𝑓𝑢 = C𝐸 Ɛ′𝑓𝑢, E𝑓 = f𝑓𝑢
Ɛ𝑓𝑢
3. Design Guidelines of Flexural Strengthening (RC Members)
Nominal strength
· Design flexural strength of a member exceeds its required factored
moment.
ɸ𝑀𝑛 ≥ 𝑀𝑢
Failure modes
· Crushing of the concrete in compression before yielding of the
reinforcing steel (Ɛ𝑐𝑢 = 0.003);
· Yielding of the steel followed by rupture of the FRP laminate (Ɛ𝑓 =
Ɛ𝑓𝑢);
· Yielding of the steel in tension followed by concrete crushing;
· Shear/tension delamination of the concrete cover;
· Debonding of the FRP from the concrete substrate;
43
ε𝑓𝑑 = 0.083√𝑓′𝑐
𝑛𝐸𝑓𝑇𝑓≤ 0.9ε𝑓𝑢 − 𝑖𝑛 𝑙𝑏 𝑢𝑛𝑖𝑡𝑠
· Note: For NSM FRP applications, the value of ε𝑓𝑑 may vary from 0.6ε𝑓𝑢
to 0.9ε𝑓𝑢 depending on member dimensions, steel and FRP
reinforcement ratios, and surface roughness of the FRP bar.
Assumptions
· Design calculations are based on the dimensions, internal reinforcing
steel arrangement, and material;
· properties of the existing member being strengthened;
· The strains in the steel reinforcement and concrete are directly
proportional to the distance from the neutral axis. That is, a plane section
before loading remains plane after loading;
· There is no relative slip between external FRP reinforcement and the
concrete;
· The shear deformation within the adhesive layer is neglected because
the adhesive layer is very thin with slight variations in its thickness;
· The maximum usable compressive strain in the concrete is 0.003;
· The tensile strength of concrete is neglected;
· The FRP reinforcement has a linear elastic stress-strain relationship to
failure.
Strain Level in FRP Reinforcement
· The effective strain level in the FRP reinforcement at the ultimate
limit state can be found from equations as follows:
ε𝑓𝑒 = ε𝑐𝑢 (d𝑓 − 𝑐
𝑐) − ε𝑏𝑖 ≤ ε𝑓𝑑
Stress Level in FRP Reinforcement
· The effective stress level in the FRP reinforcement is the maximum
level of stress that can be developed in the FRP reinforcement before
flexural failure of the section.
f𝑓𝑒 = E𝑓 ε𝑓𝑒
44
Strength reduction factor
· The use of externally bonded FRP reinforcement for flexural
strengthening will reduce the ductility of the original member.
∅ =
{
0.90 𝜀𝑡 ≥ 0.005
0.65 +0.25(𝜀𝑡 − 𝜀𝑠𝑦)
0.005 − 𝜀𝑠𝑦 𝜀𝑠𝑦 < 𝜀𝑡 < 0.005
0.65 𝜀𝑠𝑦 < 𝜀𝑡
Serviceability
· The serviceability of a member under service loads should satisfy
applicable provisions of ACI 318-05.
𝑓𝑠,𝑠 ≤ 0.80𝑓𝑦 𝑓𝑐,𝑠 ≤ 0.45𝑓𝑐′
Creep-rupture and Fatigue Stress Limits
· To avoid creep-rupture of the FRP reinforcement under sustained
stresses or failure due to cyclic stresses and fatigue of the FRP
reinforcement, the stress levels in the FRP reinforcement under these
stress conditions should be checked.
𝑓𝑓,𝑠 ≤ 𝑠𝑢𝑠𝑡𝑎𝑖𝑛𝑒𝑑 𝑝𝑙𝑢𝑠 𝑐𝑦𝑐𝑙𝑖𝑐 𝑠𝑡𝑟𝑒𝑠𝑠 𝑙𝑖𝑚𝑖𝑡
Table 2-𝑠𝑢𝑠𝑡𝑎𝑖𝑛𝑒𝑑 𝑝𝑙𝑢𝑠 𝑐𝑦𝑐𝑙𝑖𝑐 𝑠𝑡𝑟𝑒𝑠𝑠 𝑙𝑖𝑚𝑖𝑡s in FRP reinforcement
Stress type GFRP AFRP CFRP
𝑠𝑢𝑠𝑡𝑎𝑖𝑛𝑒𝑑 𝑝𝑙𝑢𝑠 𝑐𝑦𝑐𝑙𝑖𝑐 𝑠𝑡𝑟𝑒𝑠𝑠 𝑙𝑖𝑚𝑖𝑡 0.20ffu 0.30ffu 0.55ffu
Ultimate Strength of Singly Reinforced Rectangular Section
· strain level in the FRP reinforcement
ε𝑓𝑒 = ε𝑐𝑢 (d𝑓 − 𝑐
𝑐) − ε𝑏𝑖 ≤ ε𝑓𝑑
· stress level in the FRP reinforcement
f𝑓𝑒 = E𝑓 ε𝑓𝑒
· strain level in the non-prestressed steel reinforcement
ε𝑠 = (ε𝑓𝑒 + ε𝑏𝑖)ε𝑐𝑢 (𝑑 − 𝑐
d𝑓 − 𝑐)
45
· stress level in steel
f𝑠 = E𝑠 ε𝑠 ≤ f𝑦
· the assumed neutral axis depth
𝑐 =𝐴𝑠𝑓𝑠 + 𝐴𝑓𝑓𝑓𝑒𝛼1𝑓′𝑐𝛽1𝑏
· The nominal flexural strength
𝑀𝑛 = 𝐴𝑠𝑓𝑠(𝑑𝛽1𝑐
2) + 𝜑
𝑓𝐴𝑓𝑓𝑓𝑒(ℎ
𝛽1𝑐
2)
· Stress in steel under service loads
𝑓𝑠,𝑠 =[𝑀𝑠 + ε𝑏𝑖𝐴𝑓E𝑓(d𝑓 −
𝑘𝑑
3)(𝑑 − 𝑘𝑑)E𝑠
𝐴𝑠E𝑠 (d −𝑘𝑑
3) (𝑑 − 𝑘𝑑) + 𝐴𝑓E𝑓(d𝑓 −
𝑘𝑑
3)(d𝑓 − 𝑘𝑑)
· Stress in FRP under service loads
𝑓𝑓,𝑠 = 𝑓𝑠,𝑠 (E𝑓
E𝑠)
𝑑𝑓 – 𝑘𝑑
(𝑑 − 𝑘𝑑)− ε𝑏𝑖E𝑓
4. Design Guidelines of Flexural Strengthening (PC Members)
Assumptions
· Strain compatibility can be used to determine strain in the externally
bonded FRP, strain in the nonprestressed steel reinforcement, and the
strain or strain change in the prestressing steel;
· Additional flexural failure mode controlled by prestressing steel
rupture should be investigated;
· For cases where the prestressing steel is draped, several sections along
the span of the member should be evaluated
· to verify strength requirements;
· The initial strain level of the concrete substrate ε𝑏𝑖 should be calculated
and excluded from the effective strain in the FRP. The initial strain can
be determined from an elastic analysis of the existing member,
considering all loads that will be on the member at the time of FRP
46
installation. Analysis should be based on the actual condition of the
member (cracked or un-cracked section) to determine the substrate
initial strain level.
Strain Level in FRP Reinforcement
· The effective strain level in the FRP reinforcement at the ultimate
limit state for failure controlled by concrete crushing can be found
from equations as follows:
ε𝑓𝑒 = ε𝑐𝑢 (d𝑓 − 𝑐
𝑐) − ε𝑏𝑖 ≤ ε𝑓𝑑
· The effective strain level in the FRP reinforcement at the ultimate limit
state for failure controlled by prestressing steel rupture, following
equations can be used for Grade 270 and 250ksi strand, the value of
ε𝑝𝑢 is 0.035.
ε𝑓𝑒 = (ε𝑝𝑢 − ε𝑝𝑖) (d𝑓 − 𝑐
d𝑝 − 𝑐) − ε𝑏𝑖 ≤ ε𝑓𝑑
In which
ε𝑝𝑖 =p𝑒𝐴𝑝E𝑝
+p𝑒𝐴𝑐E𝑐
(1 +𝑒2
𝑟2)
Strength reduction factor
· The use of externally bonded FRP reinforcement for flexural
strengthening will reduce the ductility of the original member.
∅ =
{
0.90 𝜀𝑡 ≥ 0.005
0.65 +0.25(𝜀𝑡 − 𝜀𝑠𝑦)
0.005 − 𝜀𝑠𝑦 𝜀𝑠𝑦 < 𝜀𝑡 < 0.005
0.65 𝜀𝑠𝑦 < 𝜀𝑡
Serviceability
· The serviceability of a member under service loads should satisfy
applicable provisions of ACI 318-05.
𝑓𝑝𝑠,𝑠 ≤ 0.82𝑓𝑝𝑦 𝑓𝑝𝑠,𝑠 ≤ 0.74𝑓𝑝𝑢
47
Creep-rupture and Fatigue Stress Limits
· To avoid creep-rupture of the FRP reinforcement under sustained
stresses or failure due to cyclic stresses and fatigue of the FRP
reinforcement, the stress levels in the FRP reinforcement under these
stress conditions should be checked.
𝑓𝑓,𝑠 ≤ 𝑠𝑢𝑠𝑡𝑎𝑖𝑛𝑒𝑑 𝑝𝑙𝑢𝑠 𝑐𝑦𝑐𝑙𝑖𝑐 𝑠𝑡𝑟𝑒𝑠𝑠 𝑙𝑖𝑚𝑖𝑡
Table 2-𝑠𝑢𝑠𝑡𝑎𝑖𝑛𝑒𝑑 𝑝𝑙𝑢𝑠 𝑐𝑦𝑐𝑙𝑖𝑐 𝑠𝑡𝑟𝑒𝑠𝑠 𝑙𝑖𝑚𝑖𝑡s in FRP reinforcement
Stress type GFRP AFRP CFRP
𝑠𝑢𝑠𝑡𝑎𝑖𝑛𝑒𝑑 𝑝𝑙𝑢𝑠 𝑐𝑦𝑐𝑙𝑖𝑐 𝑠𝑡𝑟𝑒𝑠𝑠 𝑙𝑖𝑚𝑖𝑡 0.20ffu 0.30ffu 0.55ffu
Strain level in the prestressed steel
ε𝑝𝑠 = ε𝑝𝑒 +p𝑒𝐴𝑐E𝑐
(1 +𝑒2
𝑟2) + ε𝑝𝑛𝑒𝑡 ≤ 0.035
In which
ε𝑝𝑛𝑒𝑡 = 0.003(d𝑝−𝑐
𝑐) for concrete crushing failure mode
ε𝑝𝑛𝑒𝑡 = (ε𝑓𝑒 + ε𝑏𝑖) (d𝑝−𝑐
d𝑓−𝑐) for FRP rupture or debonding failure
modes
Stress level in the prestressed steel
· The stress in the prestressing steel is calculated using the material
properties of the steel. For a typical seven-wire low-relaxation
prestressing strand, the stress-strain curve may be approximated by
the following equations.
For Grade 250 ksi steel:
𝑓𝑝𝑠 = {
28500ε𝑝𝑠 𝑓𝑜𝑟 ε𝑝𝑠 ≤ 0.0076 𝑖𝑛 𝑖𝑛. −𝑙𝑏 𝑢𝑛𝑖𝑡𝑠
250 −0.04
ε𝑝𝑠 − 0.0064 𝑓𝑜𝑟 ε𝑝𝑠 > 0.0076 𝑖𝑛 𝑖𝑛. −𝑙𝑏 𝑢𝑛𝑖𝑡𝑠
For Grade 270 ksi steel:
𝑓𝑝𝑠 = {
28500ε𝑝𝑠 𝑓𝑜𝑟 ε𝑝𝑠 ≤ 0.0086 𝑖𝑛 𝑖𝑛. −𝑙𝑏 𝑢𝑛𝑖𝑡𝑠
270 −0.04
ε𝑝𝑠 − 0.007 𝑓𝑜𝑟 ε𝑝𝑠 > 0.0086 𝑖𝑛 𝑖𝑛. −𝑙𝑏 𝑢𝑛𝑖𝑡𝑠
48
· the assumed neutral axis depth
𝑐 =𝐴𝑝𝑓𝑝𝑠 + 𝐴𝑓𝑓𝑓𝑒𝛼1𝑓′𝑐𝛽1𝑏
· The nominal flexural strength
𝑀𝑛 = 𝐴𝑝𝑓𝑝𝑠(d𝑝 −𝛽1𝑐
2) + 𝜑
𝑓𝐴𝑓𝑓𝑓𝑒(d𝑓 −
𝛽1𝑐
2) 𝜑𝑓 = 0.85
· Stress in prestressing steel under service loads
ε𝑝𝑠 = ε𝑝𝑒 +p𝑒𝐴𝑐E𝑐
(1 +𝑒2
𝑟2) + ε𝑝𝑛𝑒𝑡,𝑠
ε𝑝𝑛𝑒𝑡,𝑠 =𝑀𝑠𝑒
E𝑐I𝑔 for un-cracked section at service
ε𝑝𝑛𝑒𝑡,𝑠 =𝑀𝑠𝑛𝑒𝑡𝑒
E𝑐I𝑐𝑟 for cracked section at service
· Stress in FRP under service loads
𝑓𝑓,𝑠 = (E𝑓
E𝑐)𝑀𝑠𝑦𝑏𝐼
− ε𝑏𝑖E𝑓
5. Design Guidelines of Shear Strengthening
General considerations
· For external FRP reinforcement in the form of discrete strips, the
center-to-center spacing between the strips should not exceed the sum
of d/4 plus the width of the strip.
· Completely wrapped, U-wrap and 2sides wrap are typical wrapping
schemes for shear strengthening using FRP laminates.
Table 3- reduction factors for FRP shear reinforcement 𝜑𝑓 = 0.95 Completely wrapped members
𝜑𝑓 = 0.85 Three-side and two sides
· Nominal shear strength
𝜑𝑉𝑛 ≥ 𝑉𝑢 𝜑𝑉𝑛 = 𝜑(𝑉𝑐 + 𝑉𝑠 + 𝜑𝑓𝑉𝑓)
49
𝑉𝑓 =𝐴𝑓𝑣𝑓𝑓𝑒(sin 𝛼 + cos 𝛼)𝑑𝑓𝑣
𝑠𝑓 𝐴𝑓𝑣 = 2𝑛𝑡𝑓𝑤𝑓 𝑓𝑓𝑒 = ε𝑓𝑒E𝑓
Effective strain in FRP laminates
· Completely wrapped members
ε𝑓𝑒 = 0.004 ≤ 0.75ε𝑓𝑢
· Bonded U-wraps or bonded face plies
ε𝑓𝑒 = k𝑣ε𝑓𝑢 ≤ 0.004 k𝑣 =k1k2L𝑒468ε𝑓𝑢
≤ 0.75
L𝑒 =2500
(𝑛𝑡𝑓E𝑓)^0.58 k1 = (
𝑓′𝑐
4000)
2/3
k2 =
{
𝑑𝑓𝑣 − L𝑒
𝑑𝑓𝑣 𝑓𝑜𝑟 𝑈 − 𝑤𝑟𝑎𝑝𝑠
𝑑𝑓𝑣 − L𝑒
𝑑𝑓𝑣 𝑓𝑜𝑟 2 𝑠𝑖𝑑𝑒𝑠 𝑏𝑜𝑛𝑑𝑒𝑑
· Reinforcement limits
𝑉𝑠 + 𝑉𝑓 ≤ 8√𝑓′𝑐𝑏𝑤𝑑
6. Design Guidelines of Members Subjected to Axial Force or Combined Axial and
Bending Forces
Pure axial compression
· For nonprestressed members with existing steel spiral
reinforcement
𝜑𝑃𝑛 = 0.85𝜑[0.85𝑓′𝑐𝑐( 𝐴𝑔 − 𝐴𝑠𝑡)+ 𝑓𝑦𝐴𝑠𝑡]
· For nonprestressed members with existing steel-tie
reinforcement
𝜑𝑃𝑛 = 0.8𝜑[0.85𝑓′𝑐𝑐( 𝐴𝑔 − 𝐴𝑠𝑡)+ 𝑓𝑦𝐴𝑠𝑡
50
(b) NCHRP 678: Design of FRP Systems for Strengthening Concrete Girders in Shear
Factors Affecting the Design of FRP Shear Strengthening
· Influence of FRP Properties
The effective FRP strain Ɛ𝑓𝑒 was determined based on the traditional truss analogy using the
following expression: Ɛ𝑓𝑒 = 𝑉𝑓/(𝑏𝑤𝑑𝑓E𝑓 𝜌𝑓(1 + cot 𝛽))
· Effect of Internal Transverse Steel Reinforcement
· Scale Effect
· Effect of Shear Span-to-Depth Ratio
· Influence of FRP Configuration and Anchorage
· Influence of Concrete Strength
· Influence of Fatigue
· Influence of Pre-Cracking
· Influence of Prestress
· Influence of Structural Continuity
Approaches for Relevant Changes to AASHTO LRFD Bridge Design Specifications
· The simplified procedure for evaluating Vc and Vs is given by the following equations:
𝑉𝑐 = 0.0136𝛽√𝑓′𝑐𝑏𝑣𝑑𝑣 𝑓𝑜𝑟 𝛽 = 2 𝑉𝑐 = 0.0632√𝑓′𝑐𝑏𝑣𝑑𝑣 (𝑓
′𝑐 𝑖𝑛 𝑘𝑠𝑖)
𝑜𝑟 𝑉𝑐 = 2√𝑓′𝑐𝑏𝑣𝑑𝑣 (𝑓′𝑐 𝑖𝑛 𝑝𝑠𝑖)
The contribution of the steel reinforcement is given by:
𝑉𝑠 =𝐴𝑣𝑓𝑦𝑑𝑣(cot 𝜃 + cot 𝛼) sin 𝛼
𝑠
The contribution of the FRP reinforcement is given by
𝑉𝑓 =𝐴𝑓𝑓𝑓𝑒𝑑𝑓(cot 𝜃 + cot 𝛼) sin 𝛼
𝑠𝑓
· Based on the results of statistical assessments and for simplicity, the following expressions
are proposed for determining the effective strain Ɛ𝑓𝑒
When “full-anchorage” is provided,
Ɛ𝑓𝑒 = 𝑅Ɛ𝑓𝑢 𝑤ℎ𝑒𝑟𝑒 Ɛ𝑓𝑢 =f𝑓𝑢
E𝑓 𝑅 = 4 𝜌𝑓E𝑓
−0.67 ≤ 1.0
When “full-anchorage” is not provided,
51
Ɛ𝑓𝑒 = 𝑅Ɛ𝑓𝑢 ≤ 0.012 𝑤ℎ𝑒𝑟𝑒 Ɛ𝑓𝑢 =f𝑓𝑢
E𝑓 𝑅 = 3 𝜌𝑓E𝑓
−0.67 ≤ 1.0
Maximum Spacing of Transverse Reinforcement
· If 𝑉𝑢 < 0.125𝑓′𝑐 𝑡ℎ𝑒𝑛 𝑠𝑚𝑎𝑥 = 0.8𝑑𝑣 ≤ 24.0𝑖𝑛
· If 𝑉𝑢 > 0.125𝑓′𝑐 𝑡ℎ𝑒𝑛 𝑠𝑚𝑎𝑥 = 0.4𝑑𝑣 ≤ 12.0𝑖𝑛
Nominal Shear Resistance
· The nominal shear resistance, 𝑉𝑛 , shall be determined as the lesser of:
𝑉𝑛 = 𝑉𝑐 + 𝑉𝑠 + 𝑉𝑓 + 𝑉𝑝 and 𝑉𝑛 = 0.25𝑓′𝑐𝑏𝑣𝑑𝑣 + 𝑉𝑝
In which
𝑉𝑐 = 0.0136𝛽√𝑓′𝑐𝑏𝑣𝑑𝑣
𝑉𝑠 =𝐴𝑣𝑓𝑦𝑑𝑣(cot 𝜃 + cot 𝛼) sin 𝛼
𝑠
𝑉𝑓 =𝐴𝑓𝑓𝑓𝑒𝑑𝑓(cot 𝜃 + cot 𝛼) sin 𝛼
𝑠𝑓
The effective stress of FRP shear reinforcement, f𝑓𝑒,shall be determined as:
f𝑓𝑒 = E𝑓 Ɛ𝑓𝑒, 𝑖𝑛 𝑤ℎ𝑖𝑐ℎ Ɛ𝑓𝑒 = 𝑅𝑓 Ɛ𝑓𝑢
· For completely wrapped or properly anchored U-wrap configurations
𝑅𝑓 = 0.088 ≤ 4 𝜌𝑓E𝑓−0.67 ≤ 1.0
· For Un-anchored U-wrap or Two-side bonding configurations
𝑅𝑓 = 0.066 ≤ 3 𝜌𝑓E𝑓−0.67 ≤ 1.0
Where 𝜌𝑓 shall be determined as
𝜌𝑓 =
{
2𝑛𝑓𝑡𝑓𝑤𝑓
𝑏𝑣𝑠𝑓 𝑓𝑜𝑟 𝑑𝑖𝑠𝑐𝑟𝑒𝑡𝑒 𝑡𝑟𝑖𝑝𝑠
2𝑛𝑓𝑡𝑓
𝑏𝑣 𝑓𝑜𝑟 𝑐𝑜𝑛𝑡𝑖𝑛𝑢𝑜𝑢𝑠 𝑠ℎ𝑒𝑒𝑡𝑠
Longitudinal Reinforcement
· the tensile capacity of the longitudinal reinforcement on the flexural tension side of the
member shall be proportioned to satisfy:
52
A𝑝𝑠𝑓𝑝𝑠 + A𝑠𝑓𝑦 ≥|𝑀𝑢|
𝑑𝑣ɸ𝑓+ 0.5
𝑁𝑢ɸ𝑐
+ [|𝑁𝑢ɸ𝑐
− 𝑉𝑝| − 0.5𝑉𝑠 − 0.5𝑉𝑓] cot 𝜃
𝑖𝑛 𝑤ℎ𝑖𝑐ℎ 𝑉𝑠 + 𝑉𝑓 ≤ 𝑉𝑢ɸ
· At the inside edge of the bearing area of simple end supports to the section of critical
shear, the longitudinal reinforcement on the flexural tension side of the member shall
satisfy: A𝑝𝑠𝑓𝑝𝑠 + A𝑠𝑓𝑦 ≥ (𝑉𝑢
ɸ𝑣− 0.5𝑉𝑠 − 0.5𝑉𝑓 − 𝑉𝑝) cot 𝜃
Calculation of Contribution of FRP
· The contribution of FRP 𝑉𝑓 can be computed using the 45° truss model as:
𝑉𝑓 =
{
𝐴𝑓𝑓𝑓𝑒𝑑𝑓(sin𝛼𝑓 + cos 𝛼𝑓)
𝑠𝑓
𝐴𝑓E𝑓 Ɛ𝑓𝑒𝑑𝑓(sin 𝛼𝑓 + cos 𝛼𝑓)
𝑠𝑓
𝜌𝑓E𝑓 Ɛ𝑓𝑒𝑏𝑣𝑑𝑓(sin 𝛼𝑓 + cos𝛼𝑓)
(c) NCHRP 655: Specification for the Design of Externally Bonded FRP Systems for
Repair and Strengthening of Concrete Bridge Elements
Material requirements
· The characteristic value of the glass transition temperature of the composite system shall
be at least 40° higher than the maximum design temperature 𝑇𝑚𝑎𝑥𝑑𝑒𝑠𝑖𝑔𝑛
· The characteristic value of the tensile failure strain in the direction corresponding to the
highest percentage of fibers shall not be less than 1%
· The average value and coefficient of variation of the moisture equilibrium content
determined shall not be greater than 2% and 10%, respectively. A minimum sample size of
10 shall be used in the calculation of these maximum values.
· the bond strength determined by tests specified by the Engineer of Record, shall be at least
0.064√𝑓′𝑐, where 𝑓′𝑐(𝑘𝑠𝑖) is the specified compression strength of the concrete.
Design assumptions
· The distribution of strains over the depth of the member is linear and conditions of force
equilibrium and strain compatibility are satisfied.
· Perfect bond exists between the reinforcing steel, FRP reinforcement and the concrete.
· The contribution of tension stresses in the concrete to flexural strength is neglected.
53
· The stress-strain behavior for FRP reinforcement is linear-elastic to the point of failure.
· The stress-strain behavior of steel reinforcement is bilinear, with elastic behavior up to
yielding and perfectly plastic behavior thereafter.
· The maximum usable compression strain in the concrete is equal to 0.003.
· The maximum usable strain at the FRP/concrete interface is 0.005.
· When concrete compressive strain is 0.003 under conditions of force equilibrium, it is
permitted to model the distribution of concrete stress in compression as having a uniform
stress of 0.85𝑓′𝑐 over a depth 𝑎 = 𝛽1𝑐
· When concrete compressive strain is less than 0.003 under conditions of force equilibrium,
the concrete compression stress distribution shall be modeled as parabolic according to the
following equation: f𝑐 = 2(0.9𝑓′𝑐)(
Ɛ𝑐Ɛ0)
1+(Ɛ𝑐Ɛ0)2
in which Ɛ0 = 1.71𝑓′𝑐E𝑐
Fatigue limit states
· Bridge Design Specifications, the maximum compression strain in the concrete, Ɛ𝑐 , the
strain in the steel reinforcement, Ɛ𝑠 , and the strain in the FRP reinforcement, Ɛ𝑓𝑟𝑝 , shall
meet the following requirements
Ɛ𝑐 ≤ 0.36𝑓′𝑐E𝑐 Ɛ𝑠 ≤ 0.8Ɛ𝑦 Ɛ𝑓𝑟𝑝 ≤ 𝜂Ɛ𝑢𝑓𝑟𝑝
in which
𝜂 = 0.8, 0.5, 𝑎𝑛𝑑 0.3 𝑓𝑜𝑟 𝐶𝐹𝑅𝑃, 𝐴𝐹𝑅𝑃 𝑎𝑛𝑑 𝐺𝐹𝑅𝑃
Strength limit states
· For rectangular section, factored resistance for flexure
𝑀𝑟 = 0.9[𝐴𝑠𝑓𝑠(d𝑠 − k2𝑐) + 𝐴′𝑠𝑓
′𝑠(k2𝑐 − 𝑑
′𝑠)] + 𝜑
𝑓𝑟𝑝𝑇𝑓𝑟𝑝(h − k2𝑐)
In which
𝑇𝑓𝑟𝑝 = 𝑏𝑓𝑟𝑝𝑁𝑏, k2 = 1 − 2 [(
Ɛ𝑐
Ɛ0) − arctan (
Ɛ𝑐
Ɛ0))
𝛽2 (Ɛ𝑐
Ɛ0)2 , 𝛽2 =
ln [1 + (Ɛ𝑐
Ɛ0)2
]
Ɛ𝑐
Ɛ0
Development length
· The tension development length, L𝑑, shall be taken as L𝑑 ≥𝑇𝑓𝑟𝑝
𝑏𝑓𝑟𝑝𝜏𝑖𝑛𝑡, 𝜏𝑖𝑛𝑡 = 0.065√𝑓′𝑐
Reinforcement End Peeling
54
· The peel stress at the end of externally bonded reinforcement shall meet the following
requirement 𝑓𝑝𝑒𝑒𝑙 ≤ 0.065√𝑓′𝑐 in which
𝑓𝑝𝑒𝑒𝑙 = 𝜏𝑎𝑣 [(3E𝑎E𝑓𝑟𝑝
t𝑓𝑟𝑝
t𝑎)] 𝜏𝑎𝑣 = [𝑉𝑢 + (
G𝑎E𝑓𝑟𝑝t𝑎t𝑓𝑟𝑝
)
1/2
𝑀𝑢] 𝑇𝑓𝑟𝑝(ℎ − 𝑦)
𝐼𝑇
Flexural Strengthening
· Factored resistance for flexure
𝑀𝑟 = 0.9[𝐴𝑠𝑓𝑠(d𝑠 − k2𝑐) + 𝐴′𝑠𝑓
′𝑠(k2𝑐 − 𝑑
′𝑠)] + 𝜑
𝑓𝑟𝑝𝑇𝑓𝑟𝑝(h − k2𝑐)
· For the development of the Guide Specification in this project, a nonlinear concrete model
was adopted. The stress strain relationship for such a model is defined by the following
equations: f𝑐 = 2(0.9𝑓′𝑐)(
Ɛ𝑐Ɛ0)
1+(Ɛ𝑐Ɛ0)2
in which Ɛ0 = 1.71𝑓′𝑐E𝑐
(d) Other Specifications and Guidelines
Below is a comprehensive list of specifications and guidelines pertaining to FRP
strengthening/retrofit:
ACI, 2008, Guide for the Design and Construction of Externally Bonded FRP Systems for
Strengthening Concrete Structures, ACI 440.2R-08, American Concrete Institute,
Farmington Hills, MI.
AC125, 1997, Acceptance Criteria for Concrete And Reinforced and Unreinforced
Masonry Strengthening using Fiber-Reinforced Polymer (FFP) Composite Systems , ICC
Evaluation Service, Inc., CA.
AC178, 2001, Acceptance Criteria for Inspection And Verification Of Concrete And
Reinforced And Unreinforced Masonry Strengthening Using Fiber-Reinforced Polymer
(FFP) Composite Systems, ICC Evaluation Service, Inc., CA.
CNR, 2004, Guide for the Design and Construction of Externally Bonded FRP Systems for
Strengthening Existing Structures – Materials, RC and PC structures, masonry structures,
CNR-DT 200/2004, Italian National Research Council, Rome, Italy.
DAfStb-Richtlinie: Verstärken von Betonbauteilen mit geklebter Bewehrung .
fib Bulletin No. 14, Externally bonded FRP reinforcement for RC structures , 2001, 138
pp, ISBN 2-88394-054-1.
55
fib Bulletin No. 35. Retrofitting of concrete structures by externally bonded FRPs, with
emphasis on seismic applications, 2006, 220 pp, ISBN 978-2-88394-075-8.
ISIS Design Manual No. 4 – Strengthening Reinforced Concrete Structures with
Externally—Bonded Fibre Reinforced Polymers (FRPs). ISIS Canada
JSCE, 2001, Recommendation for Upgrading of Concrete Structures with use of
Continuous Fiber Sheets, Concrete Engineering Series 41, Japan Society of Civil
Engineers, Tokyo, Japan.
SIA Norm 166, 2 Gesamtentwurf vom November 2001: Klebebewehrungen
Schweizerischer Ingenieur- und Architektenverein, Postfach, CH-8039 Zürich, 48pp
TR55, 2004, Design Guidance for Strengthening Concrete Structures Using Fibre
Composite Materials, The Concrete Society, UK.
TR57, 2003, Strengthening Concrete Structures WithFibre Composite Materials:
Acceptance, Inspection And Monitoring, The Concrete Society, UK.
2.8. Cost of FRP Wraps
Despite high material cost associated with FRP composites, the initial cost of the FRP wraps is
only a fraction of the total retrofitting cost. The remaining cost is attributed to the labor,
maintenance, and application costs (Manukonda 2011). The ease of installing, storage, handling
and transportation benefits of FRP wraps leads to a great reduction in the overall cost of the
rehabilitation. According to Lee (2005), the cost of rehabilitation was estimated at 25% of the cost
of bridge replacement. Cost effectiveness of FRPs in the rehabilitation of existing structural system
has been confirmed by many researchers (Buyukozturk and Hearing 1998, Hassan & Rizkalla
2002, Teng et al. 2007, Ilki et al. 2008, Vecchio et al. 2014). FRP composites also possess potential
lower life cycle costs (Karbhari & Zhao). The life cycle cost associated with the FRP wraps
consists of fabrication and erection cost, maintenance cost (e.g. labor, material, and equipment
cost), inspection/repair costs, and the disposal costs (Pamulaparthy 2015). Table 2.2 shows the
total cost of few FRP-retrofitted projects in the state of West Virginia (South Branch Valley
Railroad, SBVR), and other states including Wisconsin, California, Indiana, Michigan, Hawaii,
Florida, Ohio, Mississippi, Iowa, Alabama, and Cobb County Government – Georgia (Manukonda
2011). As can be seen, the cost is varied from state to state depending on many factors such as
total retrofitted areas, labor rates, material/equipment/overhead costs, etc.
56
Table 2.2 Total Contract Values of FRP-Retrofitted Projects by State DOTs (Manukonda 2011)
57
Chapter 3
FRP-Retrofitted Bridge
Projects by WVDOT
and VDOT
58
FRP composites are a promising material due to their excellent mechanical characteristics such as
high strength-to-weight ratio, corrosion free, favorable maintenance/labor costs, ease of handling
and installation, and rapid construction. They have been used in construction field to rehabilitate,
retrofit, and strengthen RC structural members for more than three decades. Major FRP bridge
applications include FRP deck/panel, FRP beam/girder, concrete deck with FRP rebar/grid, FRP
cable/tendon, FRP abutment/footing, FRP parapet/barrier/sidewalk, and FRP column/piling. West
Virginia has been recognized as a pioneer in the use of FRP composites. According to American
Composites Manufacturers Association (Busel 2016), FRP composites have been used in the
construction of approximately 220 bridges nationwide and 35 of those bridges are in West
Virginia. WVDOT Division of Highways (WVDOT-DOH) began a program to employ FRP
composites for bridge construction and rehabilitation in 1996. The first vehicular bridge with FRP
rebar reinforced concrete bridge deck in the United States (Buffalo Creek bridge; a.k.a.
McKinleyville bridge) was built in 1996 in McKinleyville, the Northern Panhandle of West
Virginia. Following the successful use of FRP composites in this bridge, many other FRP-bridge
projects (FRP deck or concrete deck reinforced by FRP rebar/grid) in the WV state were completed
during 1996-2004 period (Table 3.1). The following sections present major FRP-retrofitted bridge
projects in the state of West Virginia (Table 3.2) and few FRP-strengthened bridge projects by
Virginia Department of Transportation (VDOT). It has been found that candidate
structures/elements suitable for FRP retrofit include, but not limited to, beams/girders, slabs, bents,
columns/piles/pier caps, and abutments/footings.
Table 3.1 List of Bridges in West Virginia using FRP composites
No. Bridge Name WV
County
Year
Built/Reconstructed
or Rehabilitated
Total
Length
(ft)
Deck
Width
(ft)
Bridge
Type
(Main)
FRP
System
1 Goat Farm Jackson 2004 42.3 15 SSWB FRP deck
2 Kite Creek Monroe 2004 34.7 24.2 SSWB FRP deck
3
Howell’s Mill
(a.k.a.
“Rimmer-
White”)
Cabell 2003 237.5 32.5 CSBM FRP deck
4
Robert C.
Beach (a.k.a.
West
Buckeye)
Monongalia 2003 148.7 36.0
Timber
Arch –
Through
Type
FRP
deck/panel
59
(STTA)
5 La Chein Monroe 2003 42.8 24.0 SSWB FRP deck
6 Market Street Ohio 2001 180.5 56.0 SSPG FRP deck
7 Boy Scout
Camp Raleigh 2001 33.1 25.2 SSWB FRP deck
8 Wickwire Run Taylor 1997 34.5 21.8 SSWB FRP deck
9 Hanover Pendelton 1976/2010 118.4 28.2 SSWB FRP
deck/panel
10 Katy Truss Marion 1912/2001 90.1 13.9 SSPT FRP deck
11 Martha
Queen’s Lewis 2001 49.5 30.1 SCBB
Deck with
GFRP C-bar
and
abutment
with GFRP
rebar
12 Montrose Randolph 2001 40.7 27.5 SSWB
Deck with
FRP
rebar/grid
13 Dans Run Slab Mineral 2000 25.3 24.3 SCSL
Deck with
FRP
rebar/grid
14
Buffalo Creek
(a.k.a.
McKinleyville,
1st vehicular
bridge with
FRP rebar in
the US)
Brooke 1996 180.0 29.5 CSWB
Deck with
FRP
rebar/grid
15 North Kayford Kanawha 1940/2000 43 24.2 SCBB
Deck with
FRP
rebar/grid
16 North Acme Kanawha 1940/2001 35.3 24 SCBB
Deck with
FRP
rebar/grid
17 South Acme
(FRP Rebars) Kanawha 1940/2001 34.4 24.1 SCBB
Deck with
FRP
rebar/grid
18 Barrackville
Covered Marion 1853/1999 150 15.3 STCO
FRP
tendon/cable
19 Laurel Lick Lewis 1997 20.0 16.0 All-
composite
FRP deck,
beam, and
substructure
Note: SSWB = Simple Steel Wide-Flange Beam; SSPT = Simple Riveted Pony Truss Spans;
CSBM = Continuous Steel Stringer/Multi-Beam or Girder; STCO = Single-Span Timber
Covered Bridge; SCTB = Simple-Span Concrete T-Beams; SSPG = Structural Steel Plate Girder;
60
CSWB = Continuous Steel Wide-Flange Beam; SCBB = Simple Prestressed Concrete Box
Beam; SCSL = Single Reinforced Cast-In-Place Concrete Slab Spans;
Table 3.2 List of FRP-Retrofitted Bridges in West Virginia
No. Bridge
Name
WV
County
Year
Built/Retrofitted
Total
Length
(ft)
Deck
Width
(ft)
Bridge
Type
(Main)
FRP
System
1
Madison
Avenue
Overpass
Cabell 1966/2014 118.2 64.8 SSWB GFRP Wraps
2 East Street
Viaduct Wood 1907/2001, 2012 64.7 N/A
Concrete
slab/tunnel GFRP Wraps
3 Muddy
Creek Preston 1943/2000 129.0 29.7 SCTB CFRP Wraps
4 Flag Run Preston 1940/2002 43.2 27.0 SCTB CFRP Wraps
5
East Lynn
Lake
Campground
Wayne 1969/2014 126.5 NA NA GFRP
Jacket/Wraps
6
Pond Creek
Road
Overpass
Wood 1967/1998, 2009 NA NA NA GFRP
Jacket/Wraps
Note: SCTB = Simple-Span Concrete T-Beams; SSWB = Simple Steel Wide-Flange Beam
3.1. FRP-Retrofitted Bridge Projects by WVDOT
A total of six bridge projects retrofitted by FRP wraps between 1998 and 2014 is reported in this
section. The FRP wraps externally bonded to the concrete surface to compensate for strength lost
due to corrosion, deterioration, or fire/impact damage. The use of FRP wraps allows the
rehabilitation of the existing concrete, resulting in an economic repair as substructure replacement
generally requires replacing the entire bridge. These repairs have saved the WVDOT thousands of
dollars compared to conventional repairs. Bridge data provided herein are compiled from
inspection reports provided by the WVDOT-DOH.
Madison Avenue Overpass Bridge
Madison Avenue overpass bridge is located 0.57 miles north of Interstate I-64 in Huntington, WV
(District 2, Cabell county). This bridge was built in 1966 with 4 lanes of traffic and 16,900 average
daily traffic (as of 2012). According to WVDOT 2016 bridge inspection report, the structure
consists of three steel-beam spans (SSWB) with span lengths of 30’-0”, 57’-6” and 25’-6”
61
centerline to centerline of bearings. It is supported at both ends by reinforced concrete stub
abutments, which are founded on spread footings, and intermediately by two open-type reinforced
concrete piers. The elevation of the bottom of the footing is 567.16 at Abutment No. 1, 551.00 at
Piers No. 1 and 2, and 565.79 at Abutment No. 2. The overall length (end to end) of this bridge is
118’-2 ½”. The 7” reinforced concrete deck, which includes a ½” wearing course, is 62’-5” wide
(parapet to parapet). The asphalt wearing surface is an average 6” thick. The deck width (out to
out) is 65’-2”. WVDOT 2012 interim inspection report revealed that pier #2 was severely spalled
and delaminated. Deck and superstructure were in very good condition while piers and pier caps
are in poor condition. The geometry of the bridge was such that the two ends are at different
elevations with south end at lower elevation and north end at higher elevation. As the pier caps on
the south end were at lower elevation, they are affected severely by the rainwater seeping from
this end (Kotha 2013). Pier caps were scheduled for repair beginning March 2012 and they were
rehabilitated utilizing concrete patch and GFRP wraps. According to William (2016), total repair
cost of this bridge is approximately $47,637 ($42 per square feet) while estimated costs to replace
piers and/or entire bridge are from 1.2 to 2.5 million dollars. This indicates the cost effectiveness
of the FRP-strengthening system. It is predicted that more than one thousand bridges in the state
of West Virginia will benefit from this cost-effective repair system using FRP jackets.
62
(a) Bridge location
63
(b) Elevation view (looking west)
Spalling and exposed rebar on east cap, pier
#2
Spalling and exposed rebar on bottom of a
cap, pier #2
64
Cracking on column #1, pier #2 Cracking and delamination on pier #2
(c) Bridge condition before rehabilitation (courtesy: WVDOT 2012 interim-condition
inspection report)
(d) Applied concrete patch (top) and locations of GFRP wrap and concrete repair (bottom)
(images by Williams 2016)
65
Pier #1 Pier #1 cap
(e) Rehabilitated Piers using FRP wraps (courtesy: WVDOT 2014 periodic inspection report)
Figure 3.1 Details of Madison Avenue overpass bridge
East Lynn Lake Campground Bridge
According to Clarkson and Vijay (2015) and Vijay et al. (2016), East Lynn Lake Campground
Bridge is located on Cove Creek, which is a tributary of the East Fork of Twelve Pole Creek in
East Lynn, Wayne County, West Virginia. It was designed in 1969 following by a construction in
early 1970s as a part of a flood control project, which was the East Lynn Lake Dam. The bridge
has two lanes and five spans of continuous reinforced concrete slab structure designed for H-15-
44 loading. The bridge has the two equal end-span length of 20.25 ft. and three intermediate-span
length of 27.5 ft. each. It is supported by two abutments and four piers. Each pier is composed of
a five-pile bent with a reinforced concrete cap. All of the steel piles are founded on bedrock. The
piles are exposed to 6 ft. high water level fluctuation as lake level is lower in winter and higher in
summer. As a result, they are susceptible to continued scouring, oxidation, and rusting. H-shaped
piles of the bridge were found to be severely corroded. The U.S. Army Corps of Engineers
(USACE) reported that section losses on the steel piles were up to 60%. The corroded steel H-piles
were rehabilitated with GFRP jacket systems consisting of procured GFRP shell, self-
consolidating concrete (SCC) fill, and GFRP wrap.
66
Flag Run Bridge
Flag Run bridge is located 0.03 miles north of county route 72/6 in Preston county (District 4),
West Virginia. This reinforced concrete (RC) bridge was built in 1940 with two lanes of traffic
and 650 average daily traffic (as of 2014). It has a single span with total length of 43.2 ft. and a
span length of 40 ft. According to WVDOT 2016 bridge inspection report, the bridge
superstructure consists of four RC T-beams (33 in. high and 16.5 in. wide) topped with cast-in-
place RC slab and supported by two full-height concrete abutments. Entire bottom face and side
faces at both ends of T-beams were wrapped with CFRP composites in 2002 to achieve an HS-25
design loading. Abutments were also wrapped with CFRP and the backwalls were patched.
Elevation view End view
CFRP wraps in abutment #1 and underside of
a T-beam
CFRP wraps in abutment #2 and underside of
a T-beam
67
Overview of CFRP wraps in T-beams and an
abutment
CFRP wraps at the end of a T-beam
Figure 3.3 Details of Flag Run bridge (2016 WVDOT bridge inspection report)
Muddy Creek Bridge
Muddy Creek bridge is located 0.15 miles south of county route 26/22 in Preston county (District
4), West Virginia. This reinforced concrete (RC) bridge was built in 1943 with two lanes of traffic
and 2950 average daily traffic (as of 2015). It has three spans with a total length of 129 ft.
According to WVDOT 2017 bridge inspection report, the bridge superstructure consists of four
RC T-beams (36.75 in. high and 21 in. wide) topped with RC deck and supported by two concrete
abutments and two intermediate RC solid piers. In 2000, a CFRP wrap was added to T-beams #1
and #4 (exterior beams) to achieve an HS-25 design loading. The CFRP wraps were then covered
with a protective epoxy coating.
Elevation view (looking upstream) End view (looking north)
68
Top view of RC deck Exterior T-beams (beams #1 and #4) wrapped
with CFRP
Figure 3.4 Details of Muddy Creek bridge (2017 WVDOT bridge inspection report)
East Street Viaduct Bridge (carries CSX railroad track)
East Street Viaduct bridge is located 0.14 miles south of WV 618 in Parkersburg, West Virginia
(Wood County, District 3). This reinforced concrete (RC) bridge was built in 1907 with two traffic
lanes under and 6748 average daily traffic (as of 2015). It has a single span with a span length of
23.3 ft. and a total length of 64.7 ft. According to WVDOT 2017 bridge inspection report, the
bridge substructure is composed of unreinforced concrete full-height abutments, RC wing walls,
and bents (which are constructed of ten vertical riveted steel columns with concrete footings and
a riveted steel box beam cap). The bridge superstructure consists of 3 ft. RC top slab, which carries
2.5 ft. of railroad slag fill and ten sets of railroad tracks. In 2001, the bridge was rehabilitated with
GFRP wrapping of the abutments, wing walls, concrete bases of the bents, and top slab. In July
2012, WV State Forces, touched up the FRP on the headwall above the southbound lane with
fiberglass repair kit, applied paint to GFRP areas showing wear, and repaired the weep drains.
69
Elevation view (looking upstream) End view (looking west)
Repair of southbound lane headwall with
fiberglass repair kit
GFRP wrapping on a wing wall
Figure 3.5 Details of East Street Viaduct bridge (2012 & 2017 WVDOT bridge inspection
reports)
Pond Creek Road Overpass Bridge
Pond Creek Road Overpass bridge (a.k.a. CR 25 bridge) is located 4.26 miles north of county route
(CR) #1, which is carrying I-77 over CR #25 in Wood County, District 3, West Virginia. According
to Kotha (2013), the bridge substructure is composed of four piers and twelve 3.5 ft. circular
concrete columns. Each pier is connecting to a group of three concrete columns. In 1990, six
columns were severely damaged under fire involving a tractor trailer carrying chemical substances.
These six columns were repaired by encasing in 4 in. concrete in the same year. However, during
WVDOH routine inspection in 1994, vertical hair cracks were reported for the encased columns.
70
In 1998, three encased columns (#1, #2, #3) were repaired by hand-wrapped GFRP composite
fabric and the remaining three columns (#4, #5, #6) were strengthened by prefabricated composite
jackets. The jackets were bonded to the concrete column by filling the gap between the old column
and the jacket by epoxy grout. In 2009, three columns (#1, #2, #3) were re-hand-wrapped with
GFRP because of an another fire incident under the bridge in 2008.
Rehabilitation of six damaged columns (on left piers) (Kotha 2013)
Figure 3.6 Repair of Pond Creek bridge
In addition to the above-mentioned FRP-retrofitted bridge projects by WVDOT-DOH, there have
been several other applications of FRP in WV highway infrastructure. The South Branch Valley
Railroad in Moorefield, WV, which is owned by WVDOT, has utilized FRP wraps to repair over
50 timber piles and several bridge stringers on 100+ year old railroad bridges.
3.2. FRP-Retrofitted Bridge Projects by VDOT
Few information on FRP-retrofitted bridge projects provided by Virginia Department of
Transportation (VDOT) are given below:
1. Bridge Name: I-81 over Rte. 600 (Saumsville RD) in Shenandoah County (either VA
Structure No.: 085-2002 or 085-2003; not sure if the repair was in the northbound or
southbound bridge)
Latitude/Longitude: 38° 55' 33.8", -78° 29' 04.9"
71
Year Built: 1965
Total Length: 112 ft (3 spans, maximum span length = 37 ft)
Bridge Type: Cast-in-Place Concrete T-Beam
Deck Width: 44 ft
FRP system: wet lay-up
2. Bridge Name: Rte. 721 (Fellowship RD) over I-81 in Rockingham County (VA
Structure No. 082-6555)
Latitude/Longitude: 38° 30' 43.5", -78° 46' 56.7"
Year Built: 1965
Total Length: 300.2 ft (spans, maximum span length = 65 ft)
Bridge Type: Prestressed AASHTO Type III I-girders
Deck Width: 29 ft
FRP system: wet lay-up
3. Bridge Name: Rte. 994 (Dices Spring RD) over I-81 in Augusta County (VA
Structure No. 007-6690)
Latitude/Longitude: 38° 18' 34.5", -78° 55' 50.3"
Year Built: 1966
Total Length: 279.9 ft (5 spans, maximum span = 63 ft)
Bridge Type: Prestressed AASHTO I-girders (not sure what type)
Deck Width: 28.9 ft
FRP system: wet lay-up – NB: This strengthening was for column repair, not beam
repair
72
Chapter 4
Design Spreadsheet for
Flexural Strengthening
of RC Beams Using
FRP Composites
73
4.1. Evaluation of Concrete Structures Prior to Rehabilitation
This section presents procedures for evaluation of concrete structures prior to rehabilitation based
on “Guide for Evaluation of Concrete Structures Prior to Rehabilitation” (ACI Committee 364).
According to ACI 364.1, it is generally difficult to quantify the visible damage since it depends on
subjective criteria and the experience of the inspectors. Damage which is acceptable in one region
or one type of structure, however, may not be acceptable in another circumstance. Beginning the
field observations, some guidelines should be established in assessing the observations in order to
obtain a consistent representation and understanding of the significance of the damage. A six-point
assessment classification is recommended by ACI 364.1 as follows:
a) Unsafe
b) Potentially hazardous
c) Severe
d) Moderate
e) Minor
f) Good condition
Figure 4.1 presents a flowchart of evaluation methodology commonly undertaken before
rehabilitation. Preliminary and detailed investigations are two major tasks that containing varying
levels of some or all of the following items:
Document reviews
Field investigation
Sampling and material testing
Evaluation
Reporting
In some cases, the pertinent documentation is not available and the success of the evaluation is
dependent on the experience and judgment of the design professional.
74
Figure 4.1 Evaluation methodology (ACI 364.1R-07)
4.2. Sufficiency Rating and Overall Bridge Conditions in Virginia and West Virginia
Table 4.1 shows general condition ratings guideline for evaluating deck, superstructure, and
substructure by Federal Highway Administration (Report No. FHWA-PD-96-001 “Recording and
Coding Guide for the Structure Inventory and Appraisal of the Nation’s Bridge”). Overall bridge
conditions can be obtained by a sufficiency rating formula given in Eq. 1.
𝑆𝑢𝑓𝑓𝑖𝑐𝑖𝑒𝑛𝑐𝑦 𝑅𝑎𝑡𝑖𝑛𝑔 = 𝑆1 + 𝑆2 + 𝑆3 − 𝑆4 (1)
where S1 = Structural adequacy and safety (e.g. superstructure, substructure rating); S2 =
Serviceability and functional obsolescence (e.g. main structure type, deck condition); S3 =
Essentiality or public use (e.g. detour length, average daily traffic); and S4 = Special reductions
(e.g. traffic safety features, main structure type).
75
The result of this formula is a percentage in which 100 percent would represent an entirely
sufficient bridge and zero percent would represent an entirely insufficient or deficient bridge.
Table 4.1 General Condition Ratings for Evaluating Deck, Superstructure, and Substructure
Code Description
N NOT APPLICABLE
9 EXCELLENT CONDITION
8 VERY GOOD CONDITION - no problems noted
7 GOOD CONDITION - some minor problems
6 SATISFACTORY CONDITION - structural elements show some minor deterioration
5 FAIR CONDITION - all primary structural elements are sound but may have minor
section loss, cracking, spalling or scour
4 POOR CONDITION - advanced section loss, deterioration, spalling or scour
3 SERIOUS CONDITION - loss of section, deterioration, spalling or scour have
seriously affected primary structural components. Local failures are possible. Fatigue
cracks in steel or shear cracks in concrete may be present
2 CRITICAL CONDITION - advanced deterioration of primary structural elements.
Fatigue cracks in steel or shear cracks in concrete may be present or scour may have
removed substructure support. Unless closely monitored it may be necessary to close
the bridge until corrective action is taken
1 "IMMINENT" FAILURE CONDITION - major deterioration or section loss present in
critical structural components or obvious vertical or horizontal movement affecting
structure stability. Bridge is closed to traffic but corrective action may put back in light
service
0 FAILED CONDITION - out of service - beyond corrective action
According to 2017 National Bridge Inventory (NBI) database (FHWA 2017), West Virginia has
7,228 highway bridges and 19% of these bridges (1,372 bridges) were rated as structurally
deficient (SD). In addition, 1,394 bridges (19.3%) were rated as functional obsolete (FO). As
shown in Table 4.2, the bridge type (categorized based on bridges’ main structure type) with a
large population of SD bridges includes slab type (32.5% deficient out of 517 bridges), girder and
floorbeam system type (46.7% out of 229 bridges), tee beam type (47.1% out of 104 bridges), truss
– thru type (43.3% out of 180 bridges), arch – deck type (39.8% out of 399 bridges), channel beam
type (54.8% out of 115 bridges).
76
In the state of Virginia, a total of 13,932 highway bridges were reported by FHWA and 5.9% of
these bridges (825 bridges) was rated as structurally deficient (Table 4.3). The bridge type with
higher number of SD bridges includes truss – thru type (37.1% out of 132 bridges), girder and
floorbeam system type (25% out of 92 bridges), and tee beam type (9.6% out of 887 bridges).
Table 4.2 Bridge Conditions in West Virginia
Main Structure Type Code SD
(a)
FO
(b)
Bridge Total
(c)
a/c
(%)
b/c
(%)
Slab 01 168 144 517 32.5 27.9
Stringer/Multi-beam or Girder 02 592 516 3085 19.2 16.7
Girder and Floorbeam System 03 107 41 229 46.7 17.9
Tee Beam 04 49 25 104 47.1 24.0
Box Beam or Girders - Multiple 05 92 401 1905 4.8 21.0
Box Beam or Girders - Single or Spread 06 2 8 55 3.6 14.5
Frame (except frame culverts) 07 5 14 52 9.6 26.9
Orthotropic 08 0 2 2 0 100
Truss – Deck 09 0 4 11 0 36.4
Truss – Thru 10 78 35 180 43.3 19.4
Arch – Deck 11 159 134 399 39.8 33.6
Arch – Thru 12 1 2 8 12.5 25.0
Suspension 13 2 1 3 66.7 33.3
Stayed Girder 14 0 0 3 0 0
Movable – Lift 15 NA NA NA NA NA
Movable – Bascule 16 NA NA NA NA NA
Movable – Swing 17 NA NA NA NA NA
Tunnel 18 NA NA NA NA NA
Culvert (includes frame culverts) 19 46 36 539 8.5 6.7
Mixed types 20 1 0 1 100 0
Segmental Box Girder 21 1 1 3 33.3 33.3
Channel Beam 22 63 25 115 54.8 21.7
Other 00 6 5 17 35.3 29.4
Total 1,372 1,394 7,228 19.0 19.3
Note: SD = Structurally Deficient; FO = Functionally Obsolete
77
Table 4.3 Bridge Conditions in Virginia
Main Structure Type Code SD
(a)
FO
(b)
Bridge Total
(c)
a/c
(%)
b/c
(%)
Slab 01 57 341 1841 3.1 18.5
Stringer/Multi-beam or Girder 02 517 1602 6870 7.5 23.3
Girder and Floorbeam System 03 23 37 92 25.0 40.2
Tee Beam 04 85 276 887 9.6 31.1
Box Beam or Girders - Multiple 05 7 87 460 1.5 18.9
Box Beam or Girders - Single or Spread 06 0 4 16 0 25.0
Frame (except frame culverts) 07 4 36 218 1.8 16.5
Orthotropic 08 NA NA NA NA NA
Truss – Deck 09 2 2 7 28.6 28.6
Truss – Thru 10 49 45 132 37.1 34.1
Arch – Deck 11 10 95 219 4.6 43.4
Arch – Thru 12 0 5 10 0 50.0
Suspension 13 NA NA NA NA NA
Stayed Girder 14 0 0 1 0 0
Movable – Lift 15 0 1 4 0 25.0
Movable – Bascule 16 0 2 10 0 20.0
Movable – Swing 17 2 3 5 40.0 60.0
Tunnel 18 NA NA NA NA NA
Culvert (includes frame culverts) 19 69 84 3147 2.2 2.7
Mixed types 20 NA NA NA NA NA
Segmental Box Girder 21 0 0 4 0 0
Channel Beam 22 NA NA NA NA NA
Other 00 0 3 9 0 33.3
Total 825 2,623 13,932 5.9 18.8
Note: SD = Structurally Deficient; FO = Functionally Obsolete
4.3. Design of Flexural Strengthening of RC T-Beams Using FRP
(a) General
This section presents a design of flexural strengthening for RC Tee-beam bridge type. The use of
externally bonded FRP composites for strengthening and enhancing the load-carrying capacity of
a wide range of bridge structures has become accepted practice due to their technical and economic
benefits. Although their potential benefits have been revealed through both laboratory and field
investigations, the widespread applications of FRP systems in bridge structures has been limited
partially due to the lack of design tools. In an effort to facilitate the use of FRP materials in
strengthening reinforced concrete and prestressed bridge elements, a structural analysis
spreadsheet for the design of FRP reinforcement for flexural strengthening of reinforced concrete
78
Tee-beam bridge elements is developed. The spreadsheet will be a very useful, handy, user friendly
tool for bridge engineers to design externally bonded FRP systems for the repair and strengthening
of concrete bridges elements.
The developed spreadsheet employs the NCHRP Report 655 “Recommended Guide Specification
for the Design of Externally Bonded FRP Systems for Repair and Strengthening of Concrete
Bridge Elements” for the analysis and design of FRP laminate and all applicable provisions of the
AASHTO LRFD Bridge Design Specifications, 7th Edition (AASHTO 2014). A brief description of
inputs and design outputs is presented in subsequent subsections along with a design example to
illustrate the use of the spreadsheet and verification of results by hand calculation.
The current version of the program is limited to design FRP strengthening of reinforced concrete
Tee beams subjected to flexural loading but the spreadsheet can easily be extended for different
strengthening cases. To analyze and design a reinforced concrete tee beam retrofitted with FRP
laminates in accordance with the AASHTO LRFD Bridge Design Specifications and the NCHRP
Report 655 Recommended Guide Specification. The program analyses flexural capacity of the
existing bridge and designs externally bonded FRP laminates to increase the load carrying capacity
of the bridge. Flexural rating factors before and after strengthening are computed for design
vehicles to ensure sufficient capacity enhancement for flexure controlled bridge members.
(b) FRP Design Spreadsheet Brief Description
Bridge Geometry and Material Properties
Start to the FRP strengthening by an optional entry of the project information into the Gray cells
at the top of the first worksheet. The project information entered here becomes the header on each
page of the calculations. The spreadsheet leaves much of the formatting of the printed pages up to
the user. A basic format is already installed, but the user may add to the formatting as needed.
Generally, a BOLD BLUE cell indicates that data entry here is required for the calculations. Care
shall be exercised in entering data in the units shown adjacent to each cell to assure the spreadsheet
calculates properly.
Under the “Input” tab, basic bridge geometric data such as span length, width, number of beams,
spacing so on and so forth are required mainly for dead and live load force effects computation.
Cross section information will be utilized for section analysis of strengthen and unstrengthen
79
members. The next entries are concrete, steel, and FRP material properties. Use manufacturer’s
datasheet to enter physical properties of FRP. The rest of the first tab are various constant factors,
computed parameters based on user input data, and lookup table.
Load and Load Effects
This part of the program performs force effects due to dead and live loads the strengthened beam
will be required to resist. The spreadsheet computes live load distribution factors as per Article
4.6.2.2 and its corresponding articles of in accordance with AASHTO LRFD Bridge Design
Specifications. Skew correction factors are applied as necessary. Article and Table references to
the Specifications are highlighted in orange. Based on the NCHRP Report 655 recommended
Guide specifications for the design of externally bonded FRP systems, only Strength I and Fatigue
I limit states are considered.
Calculations
This tab begins by asking the user to select the proposed FRP strengthening techniques from a
drop-down list. The list incorporates two of the most commonly used approaches in bridge
strengthening practice.
These are either the use of Hydraulic jacking procedure to be able to perform the strengthening in
an unstressed condition or installation of FRP under stressed condition. The user has flexibility to
choose either options specific to the project at hand.
The program calculates the flexural capacity of unstrengthened beam using elastic section analysis
and estimate the amount of FRP required to satisfy strength I limit state. Estimating the depth of
the neutral axis after installation of the FRP system is an iterative process. First, an initial guess of
the depth to the neutral axis, c, has to be entered by the user. If the value entered does not satisfied
force equilibrium, use the
button to automate the iterative process. Once the button is clicked, the program will iterate until
the user input neutral axis depth converges to the true value that satisfy force equilibrium. Then
the spreadsheet calculates the steel and FRP contributions to bending resistance of the beam. From
80
the results shown here, the program alerts whether the AASHTO Strength I Load combination
limit is satisfied or not. Here, the user has a control to vary the width and number of layers of FRP
layers to meet a target flexural rating factor.
The next steps under this tab are checking ductility requirements in accordance with Article 3.4.2
of the recommended Guide Specifications and determining development length. Checking Fatigue
limit state requirements is the last step in the program conducts to make sure stress levels in
concrete, steel and FRP are within acceptable level.
Results
FRP Design spreadsheet provides two result tables. First, a summary of design checks, the final
number of FRP layers, width, and flexural capacity. Second, a summary of flexural rating factors
for design vehicles before and after FRP strengthening.
(c) Verification and Illustrative Example
The following example is presented for two purposes: (1) to illustrate how to use the developed
FRP Design spreadsheet for design of bonded FRP system for flexural strengthening of Tee
beam bridges. (2) to verify the spreadsheet results with hand calculations.
Bridge Data
Span (end -to-end) 32.5 ft
Span (CL of bearing -to-CL of bearing) 31.25 ft
Type CIP Reinforced Concrete Tee beam
Year Built 1947
Location Shenandoah county, Virginia
Concrete compressive strength, 𝑓𝑐′ 3ksi (based on year of construction)
Reinforcing steel yield strength, 𝑓𝑦 40 ksi
Girder dimensions and Steel Reinforcement See Figure 1
Shop-fabricated carbon fiber/Epoxy
composite plates
81
FRP reinforcement
Plate thickness, t=0.039"
Glass Transition Temperature: 𝑇𝑔 = 165℉
Tensile strain in the FRP reinforcement at failure: 𝜖𝑓𝑟𝑝𝑢 =
0.013
Tensile strength in the FRP reinforcement at 1% strain:
𝑃𝑓𝑟𝑝 = 9.3 𝑘𝑖𝑝𝑠/𝑖𝑛
Shear modulus of the adhesive=185 ksi
(a) Plan
82
(b) Elevation
83
(c) Typical Section
84
(d) Interior and exterior beams detail
Fig. 1 Reinforced concrete tee beam bridge FRP design example:
(a) Plan (b) Elevation (c) Section
(d) Interior and Exterior beams detail
85
(d) User Guide for FRP Design Spreadsheet for Flexural Strengthening of RC Tee Beam
Bridges
Step 1: Input Bridge geometry and material Properties data
For this illustration purpose, the interior beam is selected.
86
87
Step 2: Load and Load Effects
This tab requires only two inputs from the user, roadway width and roadway part of the
overhang.
88
Fig.2 Roadway part of the overhang, de
89
Step 3: Calculations
User has an option to select two different strengthening techniques from the drop-down list
shown below.
90
Once satisfactory design is obtained, proceed to ductility requirements check, development length
calculation and Fatigue limit state check
91
(e) Spreadsheet Results Verification
The same bridge design example used to illustrate the use of FRP design spreadsheet is used here
to verify the results. Hand calculation is employed to design FRP reinforcement for an interior
girder.
92
Step 1: Calculate Dead and Live Loads Force Effects
1.1 Dead Load Force Effects
For this illustration purpose, the barrier and curb weights are assumed to be equally shared by
interior and exterior girders.
Interior Girders
DC: Slab: (0.15) (7
12) (7.625) = 0.67 𝑘𝑖𝑝𝑠/𝑓𝑡
Girder stem: (0.15)(16)(34−7)
122 = 0.45 𝑘𝑖𝑝𝑠/𝑓𝑡
Diaphragm: 2∗(0.15)(1)(1.1667)(6.33)
31.25 = 0.08𝑘𝑖𝑝/𝑓𝑡
Curb: 2 ∗ (75𝑙𝑏
𝑓𝑡) /4 = 0.04 𝑘𝑖𝑝𝑠/𝑓𝑡
Railing and post: 2 ∗ (320𝑙𝑏
𝑓𝑡) /4 = 0.16 𝑘𝑖𝑝𝑠/𝑓𝑡
𝑤𝐷𝐶 = 1.4 𝑘𝑖𝑝𝑠/𝑓𝑡
DW: FWS
𝑤𝐷𝑊 = (0.14) (3.0
12) (7.625) = 0.27 𝑘𝑖𝑝𝑠/𝑓𝑡
𝑀𝐷𝐶 =𝑤𝐷𝐶𝐿
2
8= 170.9 𝑘𝑖𝑝 − 𝑓𝑡
𝑀𝐷𝑊 =𝑤𝐷𝑊𝐿
2
8= 32.96 𝑘𝑖𝑝 − 𝑓𝑡
𝑉𝐷𝐶 =𝑤𝐷𝐶𝐿
2= 22 𝑘𝑖𝑝𝑠
𝑉𝐷𝑊 =𝑤𝐷𝑊𝐿
2= 4.2 𝑘𝑖𝑝𝑠
93
1.2 Live Load Force Effects
1. Select Number of Lanes [A3.6.1.1.1]
NL = INT (w
12.0) = INT = (
24.0
12.0) = 2
2. Multiple Presence [A3.6.1.1.2]
# of Loaded Lanes m
1 1.2
2 1.00
3 0.85
3. Dynamic Load Allowance [A3.6.2.1]
Not applied to the design lane load.
Component IM(%)
Deck Joints 75
Fatigue 15
All other 33
4. Distribution Factors for Moment [A4.6.2.2.2]
Applicability [A4.6.2.2.1]: constant deck width, at least four parallel beams of nearly the
same stiffness, roadway part of overhang (fig…),
𝑑𝑒 = 3.29′ − 1.083′ = 2.207𝑓𝑡 < 3𝑓𝑡 OK.
94
Fig.3 Roadway part of the overhang, de
Cross section type (e) [Table A4.6.2.2.1-1]
𝑁𝑜. 𝑜𝑓 𝑏𝑒𝑎𝑚𝑠 𝑁𝑏 = 4 𝑡𝑠 = 7 𝑖𝑛.
𝑆 = 7.625𝑓𝑡 𝐿 = 31.25 𝑓𝑡
a. Interior Beams with Concrete Decks [A4.6.2.2.2b and Table A4.6.2.2.2b-1] (note that
[Table A4.6.2.2.1-2] estimate could be used as well):
(𝐾𝑔
12𝐿𝑡𝑠3)0.1 𝑜𝑟 0.25 = 1.0 𝑎𝑛𝑑
𝐼
𝐽= 1.0
One design loaded: range of applicability satisfied
𝑚𝑔𝑀𝑆𝐼 = 0.06 + (
𝑆
14)0.4
(𝑆
14)0.3
(𝐾𝑔
12𝐿𝑡𝑠3)
0.1
= 0.713
mg = girder distribution factor with multiple presence factor included
SI = single lane loaded, interior
M = moment
Two or more design lanes loaded
95
𝑚𝑔𝑀𝑀𝐼 = 0.075 + (
𝑆
9.5)0.6
(𝑆
𝐿)0.2
(𝐾𝑔
12𝐿𝑡𝑠3)
0.1
= 0.736
MI = multiple lanes loaded, interior
M = moment
For interior girders, distribution factor is governed by the multiple lanes loaded.
b. Exterior Beams [A4.6.2.2.2d and Table A4.6.2.2.2d-1]
One design lane loaded-lever rule, m=1.2 Figure 4.
Fig.4 Definition of lever rule
𝑅 = 0.5𝑃 (7.5833 + 1.583
7.5833) = 0.604𝑃
𝑔𝑀𝑆𝐸 = 0.604 SE=single lane, exterior
𝑚𝑔𝑀𝑆𝐸 = 1.2 ∗ 0.604 = 0.725
Two or more design lanes loaded, de =2.0 ft.
𝑚𝑔𝑀𝑀𝐸 = 𝑒𝑚𝑔𝑀
𝑀𝐼 ME = multiple design lanes loaded, exterior
Where 𝑒 = 0.77 +𝑑𝑒
9.1= 0.989 < 1
96
Use e=1.0
Therefore, 𝑚𝑔𝑀𝑀𝐸 = 𝑚𝑔𝑀
𝑀𝐼 = 0.736 governs
c. Distributed Live-Load Moments
Fig.5 Live-load placement for maximum bending moment: (a) Truck (b) lane (c) tandem.
Maximum Bending Moment at Midspan
Truck:
𝑀𝑇𝑟 = 32(7.81 + 0.812) + 8(0.812) = 282.4 kip-ft
Lane:
𝑀𝐿𝑛 = 0.64(7.810)(31.25)/2 = 78.1 kip-ft
Tandem:
𝑀𝑇𝑎 = 25(7.81)(1 + 11.625/15.625) = 340.52 kip-ft Governs
97
𝑀𝐿𝐿+𝐼𝑀= 𝑚𝑔𝑟 [𝑚𝑎𝑥(𝑀𝑇𝑟𝑜𝑟 𝑀𝑇𝑎) (1 +
𝐼𝑀
100) +𝑀𝐿𝑛]
Interior and exterior girders
𝑀𝐿𝐿+𝐼𝑀= 0.736(1.33 ∗ 340.52 + 78.1) = 390.81 kip-ft
Fatigue Truck Moment = (0.736)/1.2(1.15 ∗ 340.52) = 240.2 kip-ft
5. Distribution Factors for Shear [A4.6.2.2.3]
Cross section type (e) [Table A4.6.2.2.1-1], 𝑆 = 7.625𝑓𝑡, mg is independent of span
a. Interior Beams [A4.6.2.2.3a and Table A4.6.2.2.3a-1]
𝑚𝑔𝑉𝑆𝐼 = 0.36 +
𝑆
25= 0.36 +
7.625
25= 0.665
𝑚𝑔𝑉𝑀𝐼 = 0.2 +
𝑆
12− (
𝑆
35)2
= 0.2 +7.625
12− (
7.625
35)2
= 0.788, governs
b. Exterior Beams [A4.6.2.2.3b and Table A4.6.2.2.3b-1]
Lever rule 𝑚𝑔𝑉𝑆𝐸 = 0.725 governs
𝑚𝑔𝑉𝑀𝐸 = 𝑒 𝑚𝑔𝑉
𝑀𝐼
where e = 0.6 +𝑑𝑒10
= 0.80
𝑚𝑔𝑉𝑀𝐸 = 0.8 ∗ 0.788 = 0.63
c. Distributed Live-Load Shears
98
Fig.6 Live-load placement for maximum shear force: (a) truck (b) lane, and (c) tandem.
Maximum Shear Force
Truck:
𝑉𝑇𝑟 = 32(1 + 0.55) + 8(0.13) = 50.6 𝑘𝑖𝑝𝑠
Lane:
𝑉𝐿𝑛 =0.64(31.25)
2= 10 𝑘𝑖𝑝𝑠
Tandem:
𝑉𝑇𝑎 = 25 (1 +31.25 − 4
31.25) = 46.8 𝑘𝑖𝑝𝑠
Distributed Live-load Shears
𝑉𝐿𝐿+𝐼𝑀 = 𝑚𝑔𝑟 [𝑚𝑎𝑥(𝑉𝑇𝑟𝑜𝑟 𝑉𝑇𝑎) (1 +𝐼𝑀
100) + 𝑉𝐿𝑛]
For simplicity, conservatively use the maximum distribution factor i.e max(0.788 or 0.725)
Therefore, Interior and exterior girders distributed shear:
99
𝑉𝐿𝐿+𝐼𝑀 = 0.788(1.33 ∗ 50.6 + 10) = 60.9 𝑘𝑖𝑝𝑠
Fatigue Truck Shear = (0.788/1.2)(1.15 ∗ 50.6) = 38.2 kips
Summary of Maximum Moments and Shears
a) Unfactored
Load Type w (k/ft) Moment
(kip-ft)
Shear
(kip)
DC 1.4 171 22
DW 0.27 33 4
LL+IM N/A 391 61
Fatigue LL N/A 240 38
b) Factored
Limit State Moment
(kip-ft)
Shear
(kip)
Strength I 947.0 140
Service I 595 87
Fatigue I 432 69
Step 2: Calculate the flexural strength of the Tee beam
Effective depth
𝑑𝑠 = ℎ − 𝑑′
= 34 − 4.375 = 29.625𝑖𝑛.
Effective Flange Width
Effective Flange Width As per Article 4.6.2.6.1 of AASHTO LRFD Bridge Design Specifications,
the effective flange width is taken as the minimum of
· One-quarter of the effective span length;
100
· Twelve times the average depth of the slab, plus the greater of web thickness or one-half
the width of the top flange of the girder; or
· The average spacing of adjacent beams.
𝑏𝑒 = 𝑚𝑖𝑛𝑖𝑚𝑢𝑚 {
𝑙𝑒4=31.25 × 12
4= 93.75 𝑖𝑛
12𝑡𝑠 + 𝑏𝑤 = 12(7) + 16 = 100 𝑖𝑛𝑠 = 91.5 𝑖𝑛
𝑏𝑒 = 91.5 𝑖𝑛 as shown in Figure
Fig. 7 Interior beam detail
Assumptions:
· A rectangular stress block to represent the distribution of concrete compression stresses
(Article 5.7.2.2 of AASHTO LRFD Bridge Design Specifications),
· No contribution of the steel in the compression zone to the flexural strength,
· The strain in the tension steel is greater than the yield strain, and
· The neutral axis is located in the flange of the section
Thus, the compression and tension forces are 𝐶𝑐 = 0.85𝑓𝑐′𝑏𝑒𝑎 𝑎𝑛𝑑 𝑇 = 𝐴𝑠𝑓𝑦, respectively , as
illustrated in Figure 4.
From the condition of equilibrium of forces:
0.85𝑓𝑐′𝑏𝑒𝑎 = 𝐴𝑠𝑓𝑦 Thus 𝑎 =
𝐴𝑠𝑓𝑦
0.85𝑓𝑐′𝑏𝑒= 1.37 𝑖𝑛
101
Fig. 8 Force equilibrium on a reinforced concrete T-beam
The depth of the neutral axis: 𝑐 =𝑎
𝛽1=
1.37
0.85= 1.61 𝑖𝑛
Since 𝑐 < 𝑡𝑠, the assumption that depth of the neutral axis fall within the flange is appropriate.
Referring to Figure 8, the strain in the tension steel can be computed as follows:
𝜀𝑠0.003
=𝑑 − 𝑐
𝑐
𝜀𝑠 =29.6 − 1.61
1.610.003 = 0.0522
Since 𝜀𝑠 = 0.036 >𝑓𝑦
𝐸𝑠= 0.00138, the assumption that the tension steel yielded is correct.
The nominal flexural strength of the girder can then be computed from
𝑀𝑛 = 𝐴𝑠𝑓𝑦 (𝑑 −𝑎
2) = (8)(40) (29.6 −
1.61
2) = 9,214 𝑘𝑖𝑝 − 𝑖𝑛
∅𝑀𝑛 = 0.9(9214) = 8,293 𝑘𝑖𝑝 − 𝑖𝑛
Check compliance with Article 1.4.4 of the proposed Guide Specifications
𝑅𝑟 ≥ 𝜂𝑖[(𝐷𝐶 + 𝐷𝑊) + (𝐿𝐿 + 𝐼𝑀)]
∅𝑀𝑛 = 8,293 𝑘𝑖𝑝 − 𝑖𝑛 > 𝑀𝐷 +𝑀𝐿+𝐼 = 7140 𝑘𝑖𝑝 − 𝑖𝑛
Proceed with the design of an externally bonded FRP reinforcement system.
102
Step 3: Estimate the amount of FRP reinforcement required to accommodate the increase
in flexural strength.
The factored moment for Strength I limit state is
𝑀𝑢 = 11364 𝑘𝑖𝑝 − 𝑖𝑛
For a preliminary estimate of the amount of FRP reinforcement necessary to resist 947 k-ft of
moment, the following approximate design equation can be used:
𝑇𝑓𝑟𝑝 ≈𝑀𝑢 − ∅𝑀𝑛
𝑢𝑛𝑟𝑒𝑖𝑛𝑓𝑜𝑟𝑐𝑒𝑑
ℎ
𝑇𝑓𝑟𝑝 ≈(947 − 691) × 12
34= 91 𝑘𝑖𝑝𝑠
𝑇𝑓𝑟𝑝 = 𝑛𝑁𝑏𝑏𝑓𝑟𝑝
Where n is the number of FRP reinforcement plates.
Use a reinforcement width of 𝑏𝑓𝑟𝑝 = 12 𝑖𝑛 , the number of required layers is:
𝑛 =𝑇𝑓𝑟𝑝
𝑁𝑏𝑏𝑓𝑟𝑝=
91
4.65 ∗ 12= 1.63
Try 2 layers of the FRP reinforcement, for which 𝑇𝑓𝑟𝑝 = 2(4.65)(12) = 111.6 𝑘𝑖𝑝𝑠
Step 4: Compute the factored flexural resistance of the strengthened T-beam
The depth of the neutral axis can be determined from both strain compatibility and force
equilibrium
conditions as follows:
Assume c=4 in,
𝜀𝑐 =𝑐
ℎ − 𝑐𝜀𝑓𝑟𝑝 =
4
34 − 40.005 = 0.00067
𝐸𝑐 = 1820√𝑓𝑐′ = 3152 𝑘𝑠𝑖
𝜀0 = 1.71𝑓𝑐′
𝐸𝑐= 1.71
3
3152= 0.00163
103
𝜀𝑐𝜀0=0.00067
0.00163= 0.411
𝛽2 =ln [1 + (
𝜀𝑐
𝜀0)2]
𝜀𝑐
𝜀0
= 0.380
Fig.9: Reinforced Concrete T-beam externally strengthened with FRP reinforcement
Compression force in the concrete:
𝐶𝑐 = 0.9𝑓𝑐′𝛽2𝑐𝑏𝑒 = 0.9 × 3 × 0.380 × 4 × 91.5 = 375.5 𝑘𝑖𝑝𝑠
Tension Force in the tension steel:
Strain in the steel:
𝜀𝑠 =𝑑 − 𝑐
𝑐𝜀𝑐 =
29.6 − 4
40.00067 = 0.00429 > 𝜀𝑦 =
𝑓𝑦
𝐸= 0.001379
Thus, 𝑇 = 𝐴𝑠𝑓𝑦 = 8 × 40 = 320 𝑘𝑖𝑝𝑠
Tension Force in the FRP reinforcement:
𝑇𝑓𝑟𝑝 = 2(4.65)(12) = 111.6 𝑘𝑖𝑝𝑠
Total Tension Force
𝑇 = 𝑇𝑓𝑟𝑝 + 𝑇𝑠 = 111.6 + 320 = 431.6 kips
104
Clearly equilibrium of the forces is not satisfied 𝑇 − 𝐶𝑐 = 431.6 − 375.5 = 56.1 𝑘𝑖𝑝𝑠, and the
assumed depth for the neutral axis (c=4 in. ) is incorrect. By trial and error, one can find that by
assuming a depth of the neutral axis, c= 4.302 in, and repeating the above calculations, the
following values are computed:
For c=4.302 in,
𝜀𝑐 = 0.00072, 𝜀𝑠 = 0.0042 > 𝜀𝑦, 𝜀𝑐
𝜀0= 0.445, 𝛽2 = 0.406, 𝐶𝑐 = 431.53 𝑘𝑖𝑝𝑠, 𝑇𝑠 =
431.6 𝑘𝑖𝑝𝑠,
𝑇 = 𝑇𝑓𝑟𝑝 + 𝑇𝑠 = 431.6 kips, and
𝑇 − 𝐶𝑐 = 0.071 𝑘𝑖𝑝𝑠, 𝑐𝑙𝑜𝑠𝑒 𝑒𝑛𝑜𝑢𝑔ℎ 𝑡𝑜 0.
The factored flexural resistance
𝑀𝑟 = 0.9[𝐴𝑠𝑓𝑠(𝑑𝑠 − 𝑘2𝑐)] + ∅𝑓𝑟𝑝𝑇𝑓𝑟𝑝(ℎ − 𝑘2𝑐)
Where
𝑘2 = 1 −2[𝜀𝑐
𝜀0− 𝑎𝑟𝑐𝑡𝑎𝑛
𝜀𝑐
𝜀0]
𝛽2(𝜀𝑐
𝜀0)2
= 0.345
And ∅𝑓𝑟𝑝 = 0.85,
𝑀𝑟 = 0.9[8 × 40(25.6 − 0.345 × 4.302)] + 0.85 × 111.6 × (34 − 0.345 × 4.302)
= 10,030𝑘𝑖𝑝𝑠 − 𝑖𝑛
𝑀𝑟 < 11364 𝑘𝑖𝑝𝑠 − 𝑖𝑛
Either increase the width of the FRP reinforcement or the number of layers. By using three layers
of 12” wide FRP reinforcements, and re-compute the flexural resistance 𝑀𝑟. By doing so, we can
find c=4.714in. and 𝑀𝑟 = 11,505 𝑘𝑖𝑝𝑠 − 𝑖𝑛 > 11,364𝑘𝑖𝑝𝑠 − 𝑖𝑛.
Thus, AASHTO Strength I limit state is satisfied.
105
Step 5: Check ductility requirements
When reinforcing steel first yields at 𝜀𝑠 = 𝜀𝑦 =𝑓𝑦
𝐸𝑠= 0.00138. For such a case, the strain and
stress diagrams are shown in Figure 6.
Fig.10 Strain and stress distribution in the T-beam when tension steel reinforcement yield
By satisfying the conditions of force equilibrium and strain compatibility, the strain in the FRP
reinforcement when the steel tensile reinforcement yields can be found numerically to be 𝜀𝑓𝑟𝑝𝑦
=0.0016. Thus, the ductility requirement of Article 3.4.2 of the guide specification is:
𝜀𝑓𝑟𝑝𝑢
𝜀𝑓𝑟𝑝𝑦 =0.005
0.0016= 3.1 > 2.5 𝑂𝑘
Step 6: Development length
𝑙𝑑 =𝑇𝑓𝑟𝑝
𝜏𝑖𝑛𝑡𝑏𝑓𝑟𝑝=
167.4
12 × 0.065√3= 124 𝑖𝑛 = 10.33 𝑓𝑡
Distance of FRP reinforcement end termination from the girder centerline = 10.33 + 2.33 =12.66
ft. Use 13 ft and reinforce symmetrically as shown in Figure 8.
106
Fig.11 FRP Reinforcement Detail
Step 7: Check fatigue limit state
For the fatigue load combination: 0.75𝑀𝐿+𝐼 = 0.75(240) = 180 𝑘𝑖𝑝𝑠 − 𝑓𝑡 = 2160 𝑘𝑖𝑝𝑠 − 𝑖𝑛
Determine the cracking moment: 𝑀𝑐𝑟 = 𝑓𝑟𝐼𝑔
𝑦𝑡 𝑤𝑖𝑡ℎ 𝑓𝑟 = 0.24√𝑓𝑐′ = 0.416 𝑘𝑠𝑖
Section Properties:
𝐼𝑔 = 103419 𝑖𝑛4
𝑦𝑏 = 23.65 𝑖𝑛
𝑀𝑐𝑟 = 0.416103419
23.65= 1819 𝑘𝑖𝑝𝑠 − 𝑖𝑛 < 2160 𝑘𝑖𝑝𝑠 − 𝑖𝑛
Neglect the concrete part in tension and calculate the moment of inertia of an equivalent
transformed FRP section:
From the FRP reinforcement load-strain data:
𝐸𝑓𝑟𝑝 =𝑓𝑓𝑟𝑝
𝜀𝑓𝑟𝑝=𝑁𝑏 𝑡𝑓𝑟𝑝⁄
𝜀𝑓𝑟𝑝=4.65/0.039
0.005= 23850 𝑘𝑠𝑖
Modular ratio for the concrete: 𝑛𝑐 =𝐸𝑐
𝐸𝑓𝑟𝑝=
3152
23850= 0.13
Modular ratio of the steel: 𝑛𝑠 =𝐸𝑠
𝐸𝑓𝑟𝑝=
29000
23850= 1.2
Based on the modular ratios for the concrete and for the steel, an equivalent FRP transformed
section is constructed as shown below with the neutral axis assumed to lie in the flange.
107
Fig.12 Equivalent FRP transformed section
By summing the moment of areas about reference line 1-1:
𝐴𝑓𝑟𝑝 (ℎ +𝑛𝑡𝑓𝑟𝑝
2) + 𝑛𝑠𝐴𝑠𝑑 + 𝑛𝑐𝐴𝑐
𝑧
2= (𝐴𝑓𝑟𝑝 + 𝑛𝑠𝐴𝑠 + 𝑛𝑐𝐴𝑐)𝑧
𝑧 = 6.6 𝑖𝑛
Because 𝑧 = 6.6 𝑖𝑛 < 7 𝑖𝑛, the assumption that the neutral axis fall in the flange is correct.
The moment of inertia of the equivalent transformed FRP section can be computed to be
𝐼𝑡 = 9460 𝑖𝑛4
Strain in the concrete, steel reinforcement, and FRP reinforcement, respectively, due to the fatigue
load combination:
𝜀𝑐 =𝑀𝑓𝑧
𝐼𝑇𝐸𝑓𝑟𝑝=
2180 × 6.6
9460 × 23850= 0.000064 < 0.36
𝑓𝑐,
𝐸𝑐= 0.36
3
3152= 0.00034
𝜀𝑠 =𝑀𝑓(𝑑 − 𝑧)
𝐼𝑇𝐸𝑓𝑟𝑝=2180 × (29.6 − 6.6)
9460 × 23850= 0.0002 < 0.8𝜀𝑦 = 0.8
40
29000= 0.0011
𝜀𝑓𝑟𝑝 =𝑀𝑓(ℎ + 𝑡𝑓𝑟𝑝 − 𝑧)
𝐼𝑇𝐸𝑓𝑟𝑝=2180 × (34 + 3 × 0.0039 − 6.6)
9460 × 23850= 0.00026 < 𝜂𝜀𝑓𝑟𝑝
𝑢
= 0.8 ∗ 0.013 = 0.0104
Strength Limit State Load Rating Factor for Flexure
Overview
The purpose of this section is to calculate load rating factors (and/or tonnages) of the example
bridge before and after FRP flexural strengthening. This spread sheet does not provide any analysis
108
calculations, but only provides calculations for rating factors for a simple span reinforced concrete
tee beam bridge. All calculations are for the strength limit state in flexure only.
References
AASHTO 2011, The Manual for Bridge Evaluation, Second Edition
AASHTO 2014, LRFD Bridge Design Specifications, 7th Edition
Reinforced Concrete Tee Beam Flexural Capacity
Before FRP strengthening:
𝐶𝑓𝑙𝑒𝑥𝑏𝑒𝑓𝑜𝑟𝑒
= 691 𝑘𝑖𝑝 − 𝑓𝑡
After FRP strengthening:
𝐶𝑓𝑙𝑒𝑥𝑎𝑓𝑡𝑒𝑟
= 959 𝑘𝑖𝑝 − 𝑓𝑡
Dead Load Moments
DC=
DW=
Live Load Moments
Design Vehicles
𝐻𝐿93𝐿𝐿 =308 kip-ft
𝐻𝑆20𝐿𝐿 =..
Live Load Factors
Design load factors from AASHTO 2011, The manual for Bridge Evaluation, Second Edition,
Table 6A.4.2.2-1
Evaluation Level Load factor
Inventory 1.75
Operating 1.35
𝛾𝐷𝐸𝑆𝑖𝑛𝑣 = 1.75
109
𝛾𝐷𝐸𝑆𝑜𝑝𝑟 = 1.35
Rating Factor Calculations
HL-93 (Inventory)
Before
HL − 93RFinv𝑏𝑒𝑓𝑜𝑟𝑒
=∅𝑐𝐶𝑓𝑙𝑒𝑥
𝑏𝑒𝑓𝑜𝑟𝑒− 1.25𝐷𝐶 − 1.50𝐷𝑊
𝛾𝐷𝐸𝑆𝑖𝑛𝑣 × 𝐻𝐿 − 93𝐿𝐿
=1.0 ∗ 691 − 1.25(171) − 1.50(33)
1.75(308)
HL − 93RFinv𝑏𝑒𝑓𝑜𝑟𝑒
= 0.79
After
HL − 93RFinv𝑎𝑓𝑡𝑒𝑟
=∅𝑐𝐶𝑓𝑙𝑒𝑥
𝑎𝑓𝑡𝑒𝑟− 1.25𝐷𝐶 − 1.50𝐷𝑊
𝛾𝐷𝐸𝑆𝑖𝑛𝑣 × 𝐻𝐿 − 93𝐿𝐿
=1.0 ∗ 959 − 1.25(171) − 1.50(33)
1.75(308)
HL − 93RFinv𝑎𝑓𝑡𝑒𝑟
= 1.29
HL-93 (Operating)
Before
HL − 93RFopr𝑏𝑒𝑓𝑜𝑟𝑒
=∅𝑐𝐶𝑓𝑙𝑒𝑥
𝑏𝑒𝑓𝑜𝑟𝑒− 1.25𝐷𝐶 − 1.50𝐷𝑊
𝛾𝐷𝐸𝑆𝑜𝑝𝑟 ×𝐻𝐿 − 93𝐿𝐿
=1.0 ∗ 691 − 1.25(171) − 1.50(33)
1.35(308)
HL − 93RFopr𝑏𝑒𝑓𝑜𝑟𝑒
= 1.03
After
HL − 93RFopr𝑎𝑓𝑡𝑒𝑟
=∅𝑐𝐶𝑓𝑙𝑒𝑥
𝑎𝑓𝑡𝑒𝑟− 1.25𝐷𝐶 − 1.50𝐷𝑊
𝛾𝐷𝐸𝑆𝑜𝑝𝑟 ×𝐻𝐿 − 93𝐿𝐿
110
=1.0 ∗ 959 − 1.25(171) − 1.50(33)
1.35(308)
HL − 93RFopr𝑎𝑓𝑡𝑒𝑟
= 1.67
HS-20 (Inventory)
Before
HS − 20RFinv𝑏𝑒𝑓𝑜𝑟𝑒
=∅𝑐𝐶𝑓𝑙𝑒𝑥
𝑏𝑒𝑓𝑜𝑟𝑒− 1.25𝐷𝐶 − 1.50𝐷𝑊
𝛾𝐷𝐸𝑆𝑖𝑛𝑣 × 𝐻𝑆 − 20𝐿𝐿
=1.0 ∗ 691 − 1.25(171) − 1.50(33)
1.75()
HS − 20RFinv𝑏𝑒𝑓𝑜𝑟𝑒
= 0.79
After
HS − 20RFinv𝑎𝑓𝑡𝑒𝑟
=∅𝑐𝐶𝑓𝑙𝑒𝑥
𝑎𝑓𝑡𝑒𝑟− 1.25𝐷𝐶 − 1.50𝐷𝑊
𝛾𝐷𝐸𝑆𝑖𝑛𝑣 × 𝐻𝑆 − 20𝐿𝐿
=1.0 ∗ 959 − 1.25(171) − 1.50(33)
1.75()
HS − 20RFinv𝑎𝑓𝑡𝑒𝑟
= 1.291
HS-20 (Operating)
Before
HS − 20RFopr𝑏𝑒𝑓𝑜𝑟𝑒
=∅𝑐𝐶𝑓𝑙𝑒𝑥
𝑏𝑒𝑓𝑜𝑟𝑒− 1.25𝐷𝐶 − 1.50𝐷𝑊
𝛾𝐷𝐸𝑆𝑜𝑝𝑟 ×𝐻𝑆 − 20𝐿𝐿
=1.0 ∗ 691 − 1.25(171) − 1.50(33)
1.35()
111
HS − 20RFopr𝑏𝑒𝑓𝑜𝑟𝑒
= 1.03
After
HS − 20RFopr𝑎𝑓𝑡𝑒𝑟
=∅𝑐𝐶𝑓𝑙𝑒𝑥
𝑎𝑓𝑡𝑒𝑟− 1.25𝐷𝐶 − 1.50𝐷𝑊
𝛾𝐷𝐸𝑆𝑜𝑝𝑟 ×𝐻𝑆 − 20𝐿𝐿
=1.0 ∗ 959 − 1.25(171) − 1.50(33)
1.35()
HS − 20RFopr𝑎𝑓𝑡𝑒𝑟
= 1.673
Rating Factors Summary
Live Load Live Load Type Rating
Method
Inventory Flexural
Rating Factor
Operating Flexural
Rating Factor
Before After Before After
HS-20-Lane Load Only Lane LRFR
HS-20-Truck Only Axle Load LRFR
HS-20-Tandem Only Tandem LRFR
HL-93 (US) Truck + Lane LRFR 0.79 1.29 1.03 1.67
HL-93 (US) Tandem +Lane LRFR 1.29
112
Chapter 5
Project Summary and
Challenges
113
5.1. Project Summary and Challenges
The research team presents an overview of FRP wraps for rehabilitation of bridges in WVDOT
and VDOT inventory. Existing reports, national guidelines, specifications for design and
construction of externally bonded FRP are reviewed. The reviews cover all the aspects of FRP
research pertaining to repairing, reinforcing, or strengthening by external wrap and near-surface
mounting. Aspects of materials, design, construction, durability, and maintenance are covered.
National guidelines pertaining to FRP-retrofit including ACI and NCHRP are evaluated. Details
of few FRP-retrofitted bridge projects by WVDOT & VDOT are provided. Overall conditions of
all highway bridges in the state of Virginia and West Virginia are reported. These data are extracted
from the latest National Bridge Inventory by U.S. Department of Transportation, Federal Highway
Administration. FRP design spreadsheet for flexural strengthening of RC T-beams is presented.
The developed spreadsheet employs the NCHRP Report 655 “Recommended Guide Specification
for the Design of Externally Bonded FRP Systems for Repair and Strengthening of Concrete
Bridge Elements” and the AASHTO LRFD Bridge Design Specifications, 7th Edition (AASHTO
2014).
According to the most recent comprehensive survey on the use of FRP in highway
infrastructure – NCHRP Synthesis 512 (Kim 2017), FRP composites have not been widely adopted
by state DOTs and agencies yet. This is due to one or more of the following challenges: (1) Lack
of design guidelines and specifications; (2) Lack of skilled workers, designers, and contractors;
(3) Inadequate procurement procedures; (4) Limited budget; and (4) Safety concerns (e.g. risk of
fire and vandalism). Despite great efforts of many researchers in the field of FRP
strengthening/retrofitting, additional research such as long-term durability of in-situ FRP are
required to generate more technical data and to convince DOT engineers.
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