in-plane compression and shear performance of frp bridge decks acting as top chord of bridge girders

In-plane compression and shear performance of FRP bridgedecks acting as top chord of bridge girders

Thomas Keller *, Herbert Gurtler

Composite Construction Laboratory CCLab, Swiss Federal Institute of Technology, BP Ecublens, CH-1015 Lausanne, Switzerland

Available online 19 December 2004

Abstract

This paper reports on the composite action developed between pultruded FRP bridge deck systems and the supporting main gird-

ers for road bridges. Presented are requirements that an FRP deck must satisfy in order to function as part of the upper chord of the

main bridge girders. It is shown that the deck must exhibit adequate in-plane compression and shear capacity and stiffness in the

longitudinal direction of the bridge. Laboratory experiments using an existing FRP deck system with trapezoidal cell geometry

are presented and their results analyzed to establish in-plane system properties for bridge design and dimensioning. These system

properties include the effects of the material properties, the cross-sectional geometry and the adhesives used to bond the individual

pultruded shapes to form the deck. The influences of the cell geometry (trapezoidal or triangular), the fiber architecture and the

adhesive used for the deck joints (epoxy or polyurethane) on the system properties are discussed.

2004 Elsevier Ltd. All rights reserved.

Keywords: Bridges; Bridge decks; Composite action; Composite beams; Composite structures; Joints; Pultrusion

1. Introduction

Bridge decks are emerging as a promising application

for fiber-reinforced polymers (FRP). Different FRP

deck systems have already been developed and a multi-

tude of demonstration projects with smaller spans havebeen installed. In principle, two construction forms can

be discerned: deck elements from pultruded shapes that

are adhesively bonded together (e.g. ASSET, EZ-Span,

DuraSpan, Strongwell and Superdeck system) and

large-sized sandwich slabs with different core structures

(Hardcore system: stiffened foams cores or Kansas sys-

tem: thin walled cellular FRP cores). The deck elements

are usually bonded together and then fastened to pri-mary steel or concrete girders with shear studs or bolts

[1]. Adhesively bonded connections between FRP decks

and steel girders are proposed in [2] and [3]. In both

cases, the deck-to-girder connections transfer shear

forces in the longitudinal direction of the bridge between

FRP decks and girders. Depending of the degree of

shear connection stiffness, the decks act either partially

or fully as part of the top chord of the steel or concrete

girders. Such participation is known as partial or fullcomposite action [4]. Full composite action can be de-

picted by a linear axial strain distribution through the

depth of the cross-section as shown in Fig. 1, thus fulfill-

ing the hypothesis of Bernouilli [2,3]. However, as is

shown in Fig. 1, in order for a transfer of shear forces

to occur at the deck-to-girder connection, a prior shear

transfer must occur in the core of the FRP deck, from

the upper to the lower face panel (force C1 in Fig. 1).The degree of composite action, therefore, also depends

on the shear stiffness in the core-plane of the FRP deck

[3]. In order to effectively participate as part of the top

chord, FRP decks must also provide adequate in-plane

compression stiffness in the longitudinal bridge direction

between the bridge supports and adequate in-plane

0263-8223/$ - see front matter 2004 Elsevier Ltd. All rights reserved.

doi:10.1016/j.compstruct.2004.11.004

* Corresponding author.

E-mail address: [email protected] (T. Keller).

Composite Structures 72 (2006) 151–162

www.elsevier.com/locate/compstruct

mailto:[email protected]

tensile stiffness over the supports in the case of multi-

span bridges [5].

Together with the axial in-plane stiffness propertiesfor the verification of the serviceability limit state, the

axial in-plane capacity properties of the deck must also

be known for the verification of the ultimate limit state.

For single span bridges, an economical design of an

FRP–steel girder will require that the FRP deck fails

in compression before reaching the shear capacity of

the deck and the deck-to-girder connection. The com-

pression failure of the deck should occur during theyielding of the steel in the bottom part of the girder in

order to ensure ductile behavior of the hybrid girder,

as is done for steel-concrete composite bridges [6].

Keller and Gurtler determined in [3] the in-plane

compression and shear properties for a pultruded GFRP

deck system with triangular cell geometry (ASSET sys-

tem [7]). Furthermore, Keller and Schollmayer proposed

in [5] in-plane tensile properties for a pultruded GFRPdeck system with trapezoidal cell geometry (DuraSpan

766 system [8]). In both cases, property values on the

deck system level were determined. These system prop-

erties include the effects of the material properties, the

cross-sectional geometry and the adhesives used to bond

the individual pultruded shapes to form the deck. Since

these system properties are difficult to determine analyt-

ically and since bridge engineers are generally not famil-iar with fiber-reinforced polymer material science, an

experimental technique for FRP deck specimens was

developed to evaluate the structural in-plane perfor-

mance of FRP decks on the system level [3,5]. Once

the in-plane system properties are known, it will be pos-

sible to design bridge girders with compositely acting

FRP decks based on the well-known design principles

used for steel-concrete composite bridges.This paper reports on in-plane compression and shear

experiments to determine the in-plane compression and

shear system properties of a pultruded deck system with

trapezoidal cell geometry (DuraSpan 766 system). The

resulting in-plane performance of the trapezoidal cell

system will be compared to the in-plane performance

of a triangular cell system (ASSET system). As a result

of this comparison, the influences of the cross-sectional

geometry, the fiber architecture and the adhesive used

on the in-plane performance are analyzed.

2. Properties of trapezoidal cell system

The DuraSpan 766 bridge deck system consists of

pultruded shapes with a trapezoidal dual-cell cross-sec-

tion [8]. A schematic of this deck system is shown in

Fig. 2. The cross-section is composed of two face panels

and alternating vertical and inclined webs, conceived as

‘‘I-beams’’, whose flanges are connected to the deck face

panels. The deck system consists of E-glass fiber rovingsembedded in an isophthalic polyester resin. Fig. 3 (left)

shows the fiber architecture of the face panels, obtained

through resin burnout on 100 · 100 mm cutouts. The

face panel has six mats distributed over its thickness,

rovings in unidirectional (UD) and 0/±45/90 stitchedfabric form. The UD rovings are uniformly distributed

in four layers over the thickness. The total fiber content

by volume is 40% (cf. Table 1). Selected deck propertiesfrom [9] are listed in Table 2. The x-axis designates the

longitudinal direction of the bridge, while the y-axis des-

ignates the transverse bridge direction. The latter corre-

sponds to the pultrusion direction of the deck shapes.

The pultruded shapes are assembled by bonding with a

structural polyurethane adhesive (Pliogrip 6660). As

shown in Fig. 2, the adhesively bonded joints run

through the vertical webs and are stepped in the facepanels.

3. In-plane compression performance of trapezoidal cell

system

The goal of the experimental tests described below

was to determine the in-plane compression stiffnessand capacity of the trapezoidal cell system. It was

Fig. 2. Schematic of pultruded DuraSpan bridge deck system with

trapezoidal cell configuration.

Fig. 1. Principle of composite action between steel girders and FRP

bridge decks.

152 T. Keller, H. Gurtler / Composite Structures 72 (2006) 151–162

hypothesized that, if the deck were to behave in an

inelastic, non-linear manner due to progressive local

failure initiation and propagation (inter-laminar and/orthrough-thickness failures), the so-called ‘‘limit of elastic

behavior’’ would represent a further system property.

The limit of elastic behavior should not be exceeded

under the design loads in order to prevent local

delamination.

The experimental set-up for the in-plane compression

tests is shown in Fig. 4. Three specimens of 730 mm

height and 600 mm width were examined (specimens3ac–3cc). The specimens were cut and bonded by the

deck manufacturer. The specimens, arranged between

two steel plates 800 · 800 · 80 mm, were symmetrically

loaded, transverse to the pultrusion direction and along

the whole 600 mm width. Steel sections 40 · 40 mm on

both end of the specimens ensured the load introduction

and support of the deck face panels. In order to load

both deck face panels evenly, inaccuracies of the loadingsurfaces were minimized by thin lead and steel sheet

strips between the steel sections and the specimens.

The load was applied under displacement-control at a

rate of 1.5 mm/min. The Trebel press testing machine

had a load capacity of 10,000 kN.

Specimens were equipped with eight strain gages on

the face panels at mid-height to measure the strains in

the load direction and transverse to this direction onboth sides of the specimen. X-gages (PI-2-100 from To-

kyo Sokki Kenkyujo, Japan) were placed on both face

panels over the adhesively bonded joints (cf. Fig. 4).

The axial displacements over the joints could thus be

measured over a gage length of 50 mm.

Table 3 shows the measured ultimate failure loads

(average value and standard deviation) and Fig. 5 shows

the measured load–displacement values converted toaxial stress–strain curves, r–e. The axial stress was cal-

culated from the measured load divided by cross-

sectional area of the face panels (section between the

flanges of the ‘‘I-beams’’). The axial strain was calcu-

lated from the measured displacement divided by the

specimen height.

Fig. 3. Fiber architecture of the face panels (UD = unidirectional): trapezoidal (left) and triangular (right).

Table 1

Comparison of examined deck systems

Deck system Shape geometry Fiber content face panels

Cell form Depth h

[mm]

Width w

(mm)

Section area

A (mm2/m)

Deck joints Rovings

(% per vol)

Mats

(% per vol)

Total

(% per vol)

DuraSpan Trapezoidal 194 300 15,100 PU 36 4 40

ASSET [3] Triangular 225 300 15,600 Epoxy 60 2 62

Table 2

DuraSpan material properties [9] (x = bridge direction, y = transverse,

pultrusion direction)

Parameter Face panels Web walls

Ey 21.2 GPa 17.4 GPa

Ex 11.8 GPa 9.7 GPa

Ez 4.1 GPa 4.1 GPa

Gxy 5.6 GPa 7.2 GPa

Gyz, Gxz 0.6 GPa 0.6 GPa

ay 5.0 · 106/C 5.0 · 106/Cax 16.0 · 106/C 16.0 · 106/C

Fig. 4. Experimental set-up for in-plane compression tests.

T. Keller, H. Gurtler / Composite Structures 72 (2006) 151–162 153

Full load transmission from the steel plates to the

whole element width occurred after approximately0.2% of axial strain. Starting at approximately 20% of

the ultimate failure load, all specimens began to visibly

deform in a lateral direction. This lateral deformation

increased linearly with the load. At approximately 85%

of the ultimate failure load, the development of the first

delamination failures in the flanges of the ‘‘I-beams’’

was observed and reflected by a non-linear response of

the axial stress–strain curves. Subsequently, the numberof delaminated flanges and the lateral deformation

increased rapidly up to the ultimate failure load. In spec-

imen 3ac, the upper compression diagonal also buckled

in this non-linear phase. The state at the ultimate failure

and a typical local failure mode are shown in Figs. 6 and

7 (left).

The lateral deformation behavior was also observed

with the strain and X-gage measurements. Fig. 8 illus-trates the measurements of a strain and X-gage on the

left and the right side of specimen 3cc at mid-height.

The curves show an asymmetric load distribution over

the two face panels. They also show that totally concen-

tric load distribution was not obtained at the beginning

of the experiments. Using the measured strain responses

in Figs. 5 and 8, three different transverse elastic com-

pression moduli were determined in the linear-elastic

range. They are listed in Table 3 with average values

and standard deviations. The three moduli are: (a) the

elastic system modulus, calculated from the Trebel press

measurements and influenced by the lateral deformationbehavior, (b) the elastic modulus of the pultruded sec-

tions (transverse to the pultrusion direction), determined

from the strain gage readings, (c) the elastic modulus

determined from the X-gage readings over an adhesively

bonded joint.

4. In-plane shear performance of trapezoidal cell system

The goal of these experiments was to determine the

in-plane shear stiffness and capacity of the trapezoidal

cell system. The term ‘‘in-plane shear’’ is used in this

Table 3

In-plane compression and shear results for trapezoidal deck system (average values ± standard deviation)

Type of in-plane loading Ultimate

failure

load (kN)

(a) Elastic

modulus

machine (GPa)

(b) Elastic modulus

strain

gages (GPa)

(c) Elastic

modulus

X-gages (GPa)

Maximum

differential

displacement (mm)

System

shear

modulus (GPa)

Compression

(three specimens)

736 ± 19 8.1 ± 0.3 14.1 ± 2.4 6.9 ± 1.0 – –

Shear

(three specimens)

59 ± 3 – – – 33 ± 8 0.005 ± 0.001

–0.9–0.8–0.7–0.6–0.5–0.4–0.3–0.2–0.10

–40

–35

–30

–25

–20

–15

–10

–5

0

Axial Strain ε [%]

Axi

al S

tres

s σ

[MP

a]

3 ac 3 bc 3 cc idealized

–34

–29 elastic limit

Fig. 5. In-plane compression: axial stress–strain response of trapezoi-

dal specimens.

Fig. 6. In-plane compression: failure mode of trapezoidal specimen

3ac.


context on the system level and means shear in the deck

plane. Although in laminate and failure theory of FRP

materials the term ‘‘out-of-plane’’ would be used for this

type of shear, bridge designers working with traditional

materials (e.g. concrete decks) will be more familiar with

the term as it is used here. Again, the limit of elastic

behavior, in the absence of local failures, was a furtherdesign parameter to be determined. Exceeding the elastic

limit of the deck system leads to a redistribution of the

shear forces in the deck plane in the longitudinal direc-

tion of the bridge.

The experimental set-up is shown in Fig. 9. The same

general set-up was used as for the compression tests

arrangement in Fig. 4 with a modification to the load

and support conditions, which allowed for the speci-mens to be loaded in shear. Since a completely pure state

of shear could not be achieved, horizontal forces result-

ing from the eccentricity of the loading were resisted by

support angle sections seen in Fig. 9. The same specimen

geometry as in the compression experiments was used.

The three specimens examined (3as–3cs) were instru-

mented with 38 strain gages on the face panels to map

the transmission of forces between the deck face panels.

The load was applied under displacement-control at a

rate of 1.5 mm/min.

The average values and standard deviations of thefailure loads and the maximum differential displace-

ments of the face panels in the load direction are given

in Table 3. The measured load–displacement values were

transformed into a shear stress–strain curves, s–c, shownin Fig. 10. The shear load was divided by the surface of

the deck face panels to obtain the shear stress, while the

shear strain was calculated using the differential dis-

placements and the distance between the axes of thedeck face panels. In the linear-elastic range of the shear

deformation curve (c range from 0.007 to 0.015), a

system shear modulus, Gxz, of the FRP deck was deter-

mined, the average value and standard deviation of

which are given in Table 3.

During the shear test, the deck face panels remained

almost straight and parallel until ultimate failure. The

webs remained straight and no local buckling could beobserved. This can be seen in Fig. 11, which shows the

Fig. 7. In-plane compression: typical failure mode of trapezoidal (left) and triangular (right) specimen.

–1.4–1.2–1–0.8–0.6–0.4–0.200.20.4

–35

–30

–25

–20

–15

–10

–5

0

Axial Strain ε [%]

Axi

al S

tres

s σ

[MP

a]

Strain gages 3ccΩ gages 3 cc

Right Face Panel

LeftFace Panel

Fig. 8. In-plane compression: selected strain gages (on pultruded

shape) and X-gages (over bonded joint) at mid-height, trapezoidal

specimen 3cc.

Fig. 9. Experimental set-up for in-plane shear experiments.


specimen 3as subjected to the ultimate load of 63 kN.

First signs of delamination in the ‘‘I-beam’’ flanges were

observed at approximately 70% of the ultimate failure

load at the locations with maximum through-thickness

tensile stresses from combined Vierendeel and truss

action (see next paragraph). After the initiation of

delamination, the differential displacement of the deck

face panels increased non-linearly to ultimate failureload. At this load level, the specimens could be deformed

at a constant load up to approximately 7% shear strain.

Subsequently, the load decreased due to the displace-

ment-control of the experiments. The flanges of the

‘‘I-beams’’ became increasingly detached and some of

the adhesively bonded joints opened, as can be seen in

Fig. 12 (left). Joint failure, however, occurred alwaysin the adherends, never in the adhesives or in the adhe-

sive-to-adherend interfaces. The high deformability at

constant ultimate failure load is based on the system

redundancy and not on the material ductility and is des-

ignated, therefore, as system ductility.

From the strain measurement results, the axial forces

in the deck face panels and webs were determined using

the material properties listed in Table 2. The internalforces (axial and shear forces, bending moments) were

also calculated using simple structural analysis software.

Fig. 13 shows the resulting axial force, shear force and

bending moment distributions of specimen 3as at the

limit of elastic behavior (load = 43 kN). The results

from strain measurements (gage locations cf. Fig. 13)

and calculations were almost the same. From the shear

and axial force diagrams it was possible to conclude thatapproximately 85% of the load was transferred through

shear and transverse bending of the webs (Vierendeel

action). Only 15% was transferred by axial forces (truss

action) of the inclined webs. The remaining specimens

showed similar results.

5. Discussion

5.1. In-plane system properties of trapezoidal cell system

For the design of bridge girders with compositely act-

ing FRP decks, characteristic values for the in-plane

shear and compression system properties are required.

In the Eurocodes, the characteristic property values

are normally reported as 5% fractile values [10]. At thisstage in the project, fractile property values could not

yet be provided, as the number of tests performed was

too low. Average system property values are thus given

in Table 4. The values were determined for the trapezoi-

dal cell system from the experimental results shown in

Table 3 as follows:

• The average in-plane compression failure load wasconverted to an average axial failure stress, rx,u.Table 4 indicates the resulting value as well as the

stress at the limit of elastic behavior, rx,el (85% of

the axial failure stress). The elastic in-plane compres-

sion modulus, Ex, was estimated from the strain gage

and X-gage values (approximation: 2/3 from strain

gages, 1/3 from X-gages according to the correspond-

ing geometric lengths). The elastic in-plane moduluscalculated using test machine measurements was too

low due to the lateral bending that occurred in the

0 0.05 0.1 0.15 0.2 0.250

0.02

0.04

0.06

0.08

0.1

0.12

0.14

0.16

Shear Strain γ

Shea

r St

ress

τ [

MPa

]

3 as 3 bs 3 cs idealized

0.13

0.09 elastic limit

Fig. 10. In-plane shear: shear stress–strain response of trapezoidal

specimens.

Fig. 11. In-plane shear: failure mode of trapezoidal specimen 3as.


specimens. The low stiffness over the adhesive joints

reduced the in-plane stiffness of the deck considerably

(cf. Table 3). An idealized ‘‘elasto-plastic’’ axial

stress–strain curve, r–e, was derived from the mea-

sured and calculated results and is shown in Fig. 5.

The first part of the measured curves, during which

incomplete load transmission occurred, was cut.

Due to the observed system ductility, a small inelasticdeformation of 0.1% strain was considered in the ide-

alized curve. The average system property values for

the axial failure stress (34 MPa) and the elastic limit

(29 MPa) are indicated in Fig. 5.

• The in-plane shear failure load was converted to an

average shear failure stress, sxz,u. The limit of elastic

behavior and the corresponding stress, sxz,el, was

found to be 70% of the shear failure stress. An ideal-

ized ‘‘elasto-plastic’’ shear stress–strain curve, s–c,was derived from the response of the specimens

and is shown in Fig. 10. Due to the observed system

ductility, an inelastic deformation of 0.05 strain wasconsidered in the idealized curve. The average system

property values for the shear failure stress (0.13 MPa)

and the corresponding elastic limit (0.09 MPa) are

indicated in Fig. 10.

Fig. 12. In-plane shear: typical failure mode of trapezoidal (left) and triangular (right) specimen.

Fig. 13. In-plane shear: axial force, shear force and bending moment diagrams at 43 kN shear load of trapezoidal specimen 3as.

Table 4

System property values of examined deck systems (average values)

Deck system In-plane compression In-plane shear

E-Modulus

Ex (GPa)

Failure stress

rx,u (MPa)

Elastic limit

rx,el (MPa)

Failure type G-Modulus

Gxz (GPa)

Failure

stress

sxz,u (MPa)

Elastic

limit

sxz,el (MPa)

Failure

type

DuraSpan (trapezoidal) 11.7 34 29 Low system

ductility

0.005 0.13 0.09 High system

ductility

ASSET [3] (triangular) 16.2 41 None Brittle 0.047 0.61 None Brittle


5.2. Performance comparison of trapezoidal and

triangular cell system

5.2.1. Influence of cross-sectional geometry

Keller and Gurtler reported in [3] on similar experi-

ments that determined the in-plane system propertiesof the pultruded ASSET bridge deck system, composed

of bonded-together pultruded shapes with two triangu-

lar cells (cf. Fig. 15). The basic geometric dimensions,

the fiber content and the in-plane system properties of

this triangular cell system are summarized in Tables 1

and 4.

The in-plane compression properties of the two sys-

tems, given in Table 4, were similar. The idealized com-pression stress–strain behavior of the two deck systems

is shown in Fig. 16. From this comparison it was con-

cluded that the triangular deck is 38% stiffer and 21%

more resistant in compression transverse to the pultru-

sion direction than the trapezoidal deck. The trapezoidal

deck behaved in an elastic–‘‘plastic’’ manner with an

elastic limit at 85% of the compression capacity and

showed small system ductility. The triangular deck, onthe other hand, behaved linear-elastically up to brittle

failure.

In contrast to the in-plane compression behavior, the

in-plane shear behavior of the triangular and trapezoi-

dal decks was completely different. The idealized shear

stress–strain curves of the two deck systems are shown

in Fig. 17. The shear stiffness and capacity of the truss

configuration were much higher than those of the trap-ezoidal configuration: 9.4 times for the shear stiffness

and 4.6 times for the shear capacity (cf. Table 4). The

trapezoidal deck system transferred the in-plane shear

force from one face panel to the other mainly through

transverse bending of the cross-sectional web elements

(Vierendeel action, 85%), cf. Fig. 13. That is, only 15%

Fig. 14. In-plane shear: axial force, shear force and bending moment diagrams at 43 kN shear load of triangular specimen 3s.

–0.5–0.45–0.4–0.35–0.3–0.25–0.2–0.15–0.1–0.050

–45

–40

–35

–30

–25

–20

–15

–10

–5

0

Axial strain ε [%]

Axi

al s

tres

s σ

[MP

a]

DuraSpan

ASSET

–41.3

–34.1

Fig. 16. Comparison of idealized in-plane compression behavior of the

two systems.

Fig. 15. Pultruded ASSET bridge deck system with triangular cell

configuration.


of the shear force was transferred through axial forces

(truss action). Conversely, in the triangular deck system,

85% of the in-plane shear force was transferred by truss

action and 15% by Vierendeel action. Compared to the

shear force and bending moment diagrams of the trape-

zoidal deck system in Fig. 13, the triangular deck systemshowed almost no shear forces and bending moments

(cf. Fig. 14). The behavior of the trapezoidal configu-

ration showed high system ductility compared to the

triangular configuration, which exhibited very brittle

behavior. This system ductility was provided by the

redundant ‘‘frame’’ structure of the trapezoidal cross-

section, which enabled progressive local failure at

increasing load without catastrophic failure. The trian-gular structure, however, behaved more like a statically

determined truss structure and a first failure automati-

cally led to the brittle collapse of the entire structure.

5.2.2. Influence of fiber architecture

Fig. 3 (right) shows the fiber architecture of the face

panels of the triangular system, obtained through resin

burnout on 100 · 100 mm cutouts. The architecture ofthe face panels consisted of one mat on the outer side

and rovings in UD and 0/90 stitched fabric form.

The UD rovings were mainly concentrated in one center

layer. The triangular cell system had a 55% higher fiber

fraction per volume than the trapezoidal cell system (cf.

Table 1). Failure in compression of the triangular deck

system occurred in one of the truss joints at the locations

of the stepped adhesive connection, as shown in Fig. 7(right). At the onset of failure, small de-bonding cracks

in the UD-layer, parallel to the face panels, could be

observed in the exterior flanges of the stepped joints.

Subsequently, the flanges split and buckled at these loca-

tions. Failures always occurred in the adherends and

never in the adhesives or in the interfaces. Failure in

shear of the triangular deck system occurred in one of

the connections of the tension stressed diagonals: the

through-thickness tensile strength of the material was

exceeded. Fig. 12 (right) shows how the non-

anchored outer fabrics were pulled away from the joint.Failure in the trapezoidal cell system was initiated by

through-thickness tensile stresses in the layered fiber

structure for both in-plane compression and shear. At

the web-to-face panel connections the maximum

through-thickness tensile stresses due to transverse

bending in the ‘‘flanges’’ of the ‘‘I-beam-webs’’ exceeded

the through-thickness tensile strength of the material

and the ‘‘flanges’’ debonded (cf. Fig. 7 (left) and Fig.12 (left)).

From both load cases it can be concluded that failure

in both deck systems was initiated by high local

through-thickness tensile stresses at locations showing

high bending moments or axial tensile forces due to

the cross-sectional geometry. At these locations the fiber

architecture was not adapted to the type of loading.

More generally, it can be concluded that the fiber archi-tecture of both deck systems obviously was optimized

for plate bending in the direction transverse to the

bridge girders and for punching due to wheel loads,

but not for in-plane composite action loading in the lon-

gitudinal direction of the bridge. The 55% higher fiber

fraction of the triangular deck did not markedly

improve the in-plane compression behavior of the deck.

5.2.3. Influence of adhesive

The adhesively bonded joints of both deck systems

were placed at the same interval distance of 300 mm.

The joints of the triangular system were bonded with

an epoxy adhesive, while a polyurethane adhesive was

used for those of the trapezoidal system. The resulting

compression stiffness over the 50 mm gage length of

the epoxy joints (11.3 GPa [3]) was only approximately70% higher than that of the PU-joints (6.9 GPa, Table

3), even though the ratio of the E-moduli, EEP/EPU,

was on the order of 10.

The stiffness over the joints was 40% lower in the tri-

angular [3] and 51% lower in the trapezoidal cell system

as compared with average material stiffness values (from

strain gage measurements). This important stiffness

reduction due to the numerous adhesive joints will affectdirectly the efficiency of both deck systems when used as

part of the top chord of FRP–steel bridge girders.

5.3. Deck performance of FRP–steel bridge girders with

composite action

Keller and Gurtler reported in [2] and [3] on four-

point loading experiments on four full-scale bridge gird-ers (called Fix 1–4) with FRP bridge decks adhesively

0 0.01 0.02 0.03 0.04 0.05 0.06 0.07 0.08 0.090

0.1

0.2

0.3

0.4

0.5

0.6

0.7

Shear strain γ [ ]

She

ar s

tres

s τ

[MP

a]

DuraSpan

ASSET

0.61

0.13

Fig. 17. Comparison of idealized in-plane shear behavior of the two

systems.


bonded to 7.50 m span steel girders (cf. Fig. 18). The

trapezoidal DuraSpan and the triangular ASSET deck

systems were used as part of the top chords of the gird-

ers (DuraSpan girders Fix 1 and Fix 2, [2], ASSET Fix 3

and Fix 4, [3]). The deck width of the hybrid girders was

1.50 m. The adhesive connection between the FRP decksand the steel girders provided full composite action. The

behavior of the girders was predicted using the in-plane

system properties given in Table 4. In the following, the

top chord behavior of the compositely acting FRP decks

is analyzed and compared to the predictions.

Fig. 19 shows a comparison between the two systems

of the measured axial strain distributions in the mid-

span cross-section at failure, while Fig. 20 compares

the measured differential in-plane displacements be-

tween the deck face panels at one end of the two deck

systems (cf. Fig. 18 and [2,3]).

5.3.1. Top chord behavior of trapezoidal cell system

Figs. 19 and 20 show that the in-plane shear stiffness

of the trapezoidal cell system was too low to fully acti-

vate the upper face panel [2]. The measured axial strain

distribution through the depth of the cross-section of the

trapezoidal system was not linear and reached a maxi-

mum value in the lower face panel. The observed differ-

ential face panel displacements at the ends of the girders

Fig. 18. Set-up of FRP–steel bridge girders with triangular deck (deck width 1.50 m).

Fig. 19. Comparison of axial strain distributions in mid-span cross-

section at failure.

0 2 4 6 8 10 120

50

100

150

200

250

300

350

400

450

Differential shift left end [mm]

Load

per

jack

[kN

]

Fix 1, 3rd cycle DuraSpanFix 3, 3rd cycle ASSET

Fig. 20. Comparison of differential displacements (shift) between deck

face panels.


were high (10.5 mm). Consequently, failure occurred in

the lower face panel (during the yielding of the steel gir-

der). The axial stress at failure in the lower face panel,

rfail,calc, was calculated according to Eq. (1) using the ax-

ial strain measurement results on the lower face panel,

efail,meas,lo, and was found to be the same as the axial fail-ure stress, rx,u, obtained from the in-plane compression

tests:

rfail;calc ¼ efail;meas;lo Ex ¼ 0:29% 11; 700¼ 34 MPa ¼ rx;u ð1Þ

Eq. (2) shows that the in-plane shear stress in the

deck core at failure of the lower face panel, sfail,calc, cal-culated from the strain measurements on the upper face

panel, efail,meas,up, was equal to the failure stress, sxz,u,obtained from the in-plane shear tests.

sfail;calc ¼ efail;meas;up Ex tlL

¼ 0:2% 11; 700 16:83000

¼ 0:13 MPa ¼ sxz;u ð2Þ

where, t is the face panel thickness and lL the length of

shear introduction (distance of the loading points fromthe supports [2]). The differential in-plane displacement

between the deck face panels at failure Dufail,calc, was cal-culated as follows (cf. Table 4 and Fig. 17):

Dufail;calc ¼ c h ¼ sfail;calc hGxz

¼ 0:13 1945

¼ 5:0 mm ð3Þ

The measured differential in-plane displacement,

Dufail,meas, however, was 10.5 mm (cf. Fig. 20). It was

concluded that 5.5 mm of ‘‘plastic’’ deformation,

Dufail,pl, occurred at failure according to Eq. (4):

Dufail;meas ¼ Dufail;calc þ Dufail;pl ¼ 5:0þ 5:5 ¼ 10:5 mm

ð4ÞThe differential in-plane displacement between the

deck face panels at serviceability (SLS, 80 kN load per

loading point [2]), DuSLS,calc, calculated from the strain

measurements on the upper face panel at SLS, eSLS,meas,up,(0.035% in [2]), was almost the same as the measured

value (0.8 mm at 80 kN, cf. Fig. 20):

DuSLS;calc ¼eSLS;meas;up Ex t h

lL Gxz

¼ 0:035% 11; 700 16:8 1943000 5 ¼ 0:9 mm

ffi DuSLS;meas ¼ 0:8 mm ð5Þ

5.3.2. Top chord behavior of triangular cell system

Figs. 19 and 20 show that the in-plane shear stiffness

of the triangular cell system almost fully activated the

upper face panel [3]. Only a small change in slope in

the linear axial strain distribution was observed and

the differential face panel displacements were small

(0.3 mm). Consequently, failure occurred in the upper

face panel (during the yielding of the bottom steel

flange). The axial stress at failure in the upper face pa-

nel, rfail,calc, was calculated according to Eq. (6) using

the strain measurement results on the upper face panel,

efail,meas,up:

rfail;calc ¼ efail;meas;up Ex ¼ 0:20% 16; 200¼ 32 MPa < rx;u ¼ 41 MPa ð6Þ

The resulting stress at failure was 22% below the fail-

ure stress, rx,u, obtained from the in-plane compression

tests (Table 4). The difference could be explained by a

premature deck failure caused by the experimental set-up, cf. [3].

The in-plane shear stress in the deck core at failure of

the upper face panel, sfail,calc, calculated from the strain

measurement results on the upper face panel, efail,meas,up,

was far below the failure stress, sxz,u, obtained from the

in-plane shear tests:

sfail;calc ¼ efail;meas;up Ex tlL

¼ 0:2% 16; 200 15:63000

¼ 0:17 MPa < sxz;u ¼ 0:61 MPa ð7Þ

The differential in-plane displacement betweenthe deck face panels at failure of the upper face panel,

Dufail,calc, was calculated according to Eq. (8) and over-

estimated the measured value (Fig. 20). However, both

measured and calculated values were very small and

within the margin of error that can be expected in a

full-scale experiment.

Dufail;calc ¼ c h ¼ sfail;calc hGxz

¼ 0:17 22547

¼ 0:8 mm

> 0:3 mm ðmeasuredÞ ð8Þ

5.4. Standardization of the system experiments

The results presented in the previous section led to

the conclusion that the proposed deck property values

on the system level could enable sufficiently accurate

modeling of the behavior of both FRP deck types acting

as part of the top chord of FRP–steel bridge girders.Furthermore, as described in [2] and [3], the deflection

behavior and the failure loads of the FRP–steel girders

could also be predicted using these system properties.

The proposed experiments to determine the in-plane sys-

tem properties of FRP decks, therefore, could be stan-

dardized and also used for other FRP deck types.

6. Conclusions

Based on the results of the investigations presented,

the following conclusions can be made:


(1) An experimental technique was proposed to evalu-

ate the structural in-plane performance of FRP

decks acting as part of the top chord of bridge gird-

ers. These system properties include the effects of

the material properties, the cross-sectional geome-

try and the adhesives used for the deck joints.Compression and shear tests were performed on

specimens of the pultruded DuraSpan 766 bridge

deck with trapezoidal cell geometry. The resulting

in-plane system properties were compared to those

obtained from similar tests on the pultruded

ASSET bridge deck system with triangular cell

geometry. The cell geometry, fiber architecture

and type of adhesive influenced the in-plane systemproperties.

(2) The in-plane compression behavior of the two deck

systems was similar. The compression stiffness and

capacity transverse to the pultrusion direction

reflected similar values. The adhesively bonded

joints between the single pultruded profiles affected

considerably the in-plane compression stiffness and,

therefore, the efficiency of the decks to act as thetop chord of bridge girders.

(3) The in-plane shear behavior of the two deck sys-

tems differed significantly. The shear stiffness and

capacity depended strongly on the geometry of

the cell structure. The triangular configuration

transferred the shear forces from one face panel

to the other mainly by axial forces (truss action)

and was much stiffer and resistant than the trape-zoidal configuration, which transferred the shear

forces mainly by transverse bending (Vierendeel

action). The triangular configuration with an

in-plane system shear modulus of 0.05 GPa enabled

almost full composite action over the depth of

full-scale FRP–steel bridge girders. Conversely,

with the trapezoidal configuration, which showed

an approximately 10 times lower system shearmodulus, only partial composite action could be

achieved.

(4) The failure mode of the triangular configuration

was abrupt and brittle. The trapezoidal configura-

tion, however, showed high system ductility due

to its redundant structural system, with progressive

local failure initiation and propagation. For both

configurations, failures always occurred in the adh-erends and never in the adhesives or in the adher-

end-to-adhesive interfaces. In both cases the fiber

architecture was not optimized for in-plane loading

from composite action.

(5) The proposed deck property values on the system

level enabled sufficiently accurate prediction of the

behavior of FRP decks acting as part of the topchord of bridge girders. The in-plane system prop-

erty values allowed for the use of standard design

methods for bridge girders with compositely acting

FRP decks. The proposed experiments to determine

the in-plane system properties of FRP decks could

be standardized and also used for other FRP deck

systems.

Acknowledgments

The authors wish to acknowledge the support of the

Swiss Federal Roads Authority, Martin Marietta Com-

posites, Raleigh USA (supplier of the DuraSpan 766bridge deck specimens) and Fiberline A/S, Denmark

(supplier of the ASSET bridge deck specimens).

References

[1] Keller T. Use of fibre reinforced polymers in bridge construction.

Structural Engineering Documents 7. Int Assoc Bridge Struct

Eng. ISBN 3-85748-108-0, 2003: 131p.

[2] Keller T, Gurtler H. Composite action and adhesive bond

between FRP bridge decks and main girders. J Compos Constr,

in press.

[3] Keller T, Gurtler H. Quasi-static and fatigue performance of a

cellular FRP bridge deck adhesively bonded to steel girders.

Compos Struct, in press.

[4] Cassity P, Richards D, Gillespie J. Composite acting FRP deck

and girder system. Struct Eng Int 2002;12(2):71–5.

[5] Keller T, Schollmayer M. In-plane tensile performance of a

cellular FRP bridge deck acting as top chord of continuous bridge

girders. Compos Struct, in press.

[6] Eurocode 4. Design of composite steel and concrete structures.

Part 2: Rules for bridges. European Committee for Standardiza-

tion, Brussels, prEN 1994-2, 2002.

[7] ASSET bridge deck system: www.fiberline.com.

[8] DuraSpan bridge deck system: www.martinmariettacomposites.

com.

[9] DARPA Final Technical Report. Advanced composites for bridge

infrastructure renewal––Phase II, Task 16. Modular composite

bridge, 2002.

[10] Eurocode. Basis of structural design. European Committee for

Standardization, Brussels, EN 1990, 2002.


http://www.fiberline.com

http://www.martinmariettacomposites.com

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