multifunctional all-gfrp joint for concrete slab structures

Construction

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Construction and Building Materials 21 (2007) 1206–1217

and Building

MATERIALS

Multifunctional all-GFRP joint for concrete slab structures

Thomas Keller *, Florian Riebel, Aixi Zhou

Composite Construction Laboratory (CCLab), Swiss Federal Institute of Technology Lausanne (EPFL), BAT BP, Station 16,

CH-1015 Lausanne, Switzerland

Received 28 October 2005; received in revised form 16 May 2006; accepted 15 June 2006Available online 17 August 2006

Abstract

An existing hybrid-GFRP/steel joint for load transfer and thermal insulation in concrete slab structures was developed into an all-GFRP joint. The new joint consists of a pultruded GFRP tensile/shear element anchored through adhesively bonded ribs in the concreteand a compression/shear element with contacting cap-plates. The new joint improves considerably the energy savings of buildings due tothe low thermal conductivity of GFRP composites. The quasi-static performance of the new joint was investigated through full-scaleexperiments on cantilever beam elements and analytical modeling. The all-GFRP joint provides ductile performance similar to thehybrid-GFRP/steel joint. The transfer of bending moments occurs through tensile forces in the upper flange of the tensile/shear elementand compression forces in the lower flange of the compression/shear element. The shear force transfer is shared between both GFRPelements. The tensile/shear element attracts 50% more shear force due to the fixed end supports. Simple analytical models can be usedto describe the joint behavior and compare well with measurements.� 2006 Elsevier Ltd. All rights reserved.

Keywords: Concrete slabs; Fiber composites; Multiple purpose structures; Pultrusion; Thermal insulation

1. Introduction

In building construction today, a significant develop-ment towards energy saving construction concepts andmethods can be observed. New codes with stringentrequirements regarding thermal insulation have beenimplemented, e.g. the new Swiss Code SIA 380/1 [1], withdecreased values for linear thermal bridge allowance limitsin facade constructions. The new limiting value of0.3 W/m K is approximately half of former limiting val-ues, which were not specified exactly in the former codes.For this reason, thermal bridges in insulating facades dueto penetrations of structural components (e.g. cantileverbalcony structures) and built from materials with highthermal conductivity (concrete or steel), are a major con-cern [2]. Based on the new requirements, efforts are beingmade by the construction industry to develop new struc-tural components with improved thermal behavior. An

0950-0618/$ - see front matter � 2006 Elsevier Ltd. All rights reserved.

doi:10.1016/j.conbuildmat.2006.06.003

* Corresponding author. Tel.: +41 21 693 3226; fax: +41 21 693 6240.E-mail address: [email protected] (T. Keller).

example of this was presented by Keller et al. [2] whodeveloped a hybrid-GFRP/steel joint for thermal insula-tion and load transfer in concrete slabs. The joint, shownin Fig. 1, is inserted in cantilever slabs at the location ofthe facade penetration, providing a transfer of shearforces and bending moments as well as thermal insulation.The use of new high strength GFRP materials (GlassFiber-Reinforced Polymers) with very low thermal con-ductivity allowed for a reduction of approximately 70%in the linear thermal bridge allowance. Although thenew joint is characterized by higher material costs, ithas proven to be competitive due to the multifunctionalityoffered by the relatively expensive GFRP materials.

The present paper describes the further development ofthe hybrid-GFRP/steel joint presented in Ref. [2] to an all-GFRP joint, shown in Fig. 2. The all-GFRP joint has asimilar load transfer capacity and provides a further reduc-tion of approximately 50% in the linear thermal bridgeallowance (from 0.1–0.2 W/m K to 0.05–0.10 W/m K).Compared with the all-steel joint used in earlier years (lin-ear thermal bridge allowance of 0.4–0.6 W/m K [3]), an

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Fig. 2. New thermally insulating all-GFRP joint in balcony slab atinsulating layer location of facade.

Fig. 1. Existing hybrid-GFRP/steel joint.

T. Keller et al. / Construction and Building Materials 21 (2007) 1206–1217 1207

overall reduction of almost 90% of the linear thermalbridge allowance could be achieved, while maintainingalmost the same load transfer capacity. To verify theload-bearing behavior of the new all-GFRP joint, quasi-static full-scale experiments on concrete cantilever beamswith integrated insulating joints were performed. Consider-ing the configuration of the joints being tested, it wasassumed that the behavior of the beams with all-GFRPjoints can be used directly to determine the behavior of

Fig. 3. CS-element cross-section with cap-plates (left), T

slabs. The durability of the GFRP elements in alkalinesolutions at different temperatures was also investigatedover a period of 18 months and led to successful results.Only the results of the quasi-static experiments on beamswith all-GFRP joints, however, are presented in the follow-ing and compared to the results of the beams with hybrid-GFRP joints.

2. All-GFRP joint with insulating GFRP elements

The design of the new all-GFRP joint was based on thepreviously developed hybrid-GFRP/steel joint, which con-sisted of a pultruded GFRP three-cell compression/shearelement (CS-element) with cap-plates on each side and steelbars for the transfer of shear and upper tensile forces due tobending moments [2]. The GFRP CS-element, which trans-fers the lower compression forces due to negative bendingmoments and a part of the shear force from the cantilever-ing concrete slab, was retained unchanged in the new joint.The investigation of the hybrid-GFRP/steel joint behaviorshowed that between 43% and 63% of the shear forces aretransferred by the GFRP element webs in a shear dominantloading mode through element tilting. In a moment domi-nated loading mode with smaller shear forces, however,no shear force is transferred by the element.

In the new joint, the remaining steel bars of the hybridjoint were replaced by a new combined GFRP tensile/shearelement (TS-element), as shown in Fig. 2. The 100 · 60 mmcross-section of the TS-elements, shown in Fig. 3, is identi-cal to that of the CS-elements. They are cut from the samepultruded profiles as the CS-elements, but placed upsidedown in the joint, with the thicker 16 mm flange on theupper side to maximize the inner lever arm between thelower compression and upper tensile flange, see Fig. 3.The four webs of the three cells are 5 mm thick each, whilethe lower flange is 3 mm deep. Both elements, TS and CS,are placed at a distance of 150 mm or 300 mm in the insu-lating joint of the concrete slab, and are embedded into sty-rofoam blocks.

S-element cross-section with anchorage ribs (right).

Fig. 4. Side and top view of TS-elements with strain gage arrangement,two ribs (above) and three-ribs (below, supported side).

Fig. 6. Anchorage part of a three-rib element installed in formwork ofsupported beam side (top view).

1208 T. Keller et al. / Construction and Building Materials 21 (2007) 1206–1217

The transfer of the tensile forces from the TS-element tothe concrete is more difficult than for the CS-element,which transfers the forces by simple contact pressure andstatic friction through the cap-plates [2]. The TS-elementmust be anchored into the concrete. Two anchorage vari-ants were investigated within this project: both variantsextended the element into the concrete and used ribs adhe-sively bonded onto the upper flange and the sides of theTS-element, as shown in Figs. 3 and 4. The first anchoragevariant consisted of two ribs 10 mm in depth and 60 mm inlength on each side of the GFRP joint, while the secondvariant consisted of three ribs of the same depth and80 mm length also on both sides. The ribs were cut frompultruded GFRP flat sections. The distance between theconcrete edge and the first rib from the joint was kept con-

Fig. 5. Typical experimental beam (M2) wit

stant, 100 mm, to ensure the introduction of the anchorageforce of the first rib into the concrete. The distance betweenthe ribs was 50 mm to provide the space for steel shear stir-rups, as shown in Figs. 5 and 6. The total length of the two-rib TS-element was 680 mm and 1020 mm for the three-ribTS-element. An anchorage failure was expected for theshorter anchorage length, while no anchorage failure wasaimed for the longer length. No failure in the adhesivebond was expected at a rib length of 60 mm. This bondlength resulted from an adhesive joint design accordingto Keller and Vallee [4], assuming that each upper ribanchors the same portion of the total tensile force fromthe bending moment and neglecting the participation ofthe lateral ribs.

The materials used for the GFRP elements are isophtha-lic polyester resin and 60% glass fibers by weight (80% UD-rovings and 20% combined mats stitched together fromfabrics and CSM). The compressive strength and elasticmodulus of the CS-elements with bonded cap-plates wereobtained from concentrical compression tests. Table 1 sum-marizes the test results from 14 specimens [2]. Similarly,axial tensile tests were performed on the 16 mm flangescut from the TS-elements. Table 1 also summarizes the testresults from 3 specimens. According to the manufacturer’sdesign manual [5], the tensile and compressive strength is240 MPa while the shear strength is 25 MPa. These valuesare conservative compared with the measured results, par-ticularly the tensile strength.

h new joint and concrete reinforcement.

Table 1Properties of CS-element (14 tests) and TS-element (3 tests), average values ± standard deviation

Element Ultimate failure load Fu (kN) Strength fu (MPa) Elastic modulus E (GPa)

Compression/shear CS (system with cap-plates) �727 ± 64 �265 ± 23 (compression) 16.0 ± 0.5Tensile/shear TS (16 mm flange) 861 ± 23 538 ± 15 (tension) 40.5 ± 0.6


The ribs of the TS-elements are bonded to the elementswith a two-component epoxy adhesive (SikaDur 330). Thechosen adhesive thickness was 2 mm due to the wires of thestrain gages embedded in the adhesive (see next section).To ensure a constant thickness of the adhesive layer, threehigh precision glass bead spacers were placed into theliquid epoxy resin during joint manufacturing. The surfacessubjected to bonding were prepared as follows: cleaningand degreasing using acetone, mechanical abrading of thesurface veil with a sander until to the first visible mat fibersand subsequent re-cleaning and re-degreasing of the sur-faces using acetone.

A main concern of the hybrid-FRP and all-FRP jointswas the long-term performance of the pultruded GFRP ele-ments and adhesive bonds exposed to concrete pore watersolution. Therefore, a durability study was performed onthe CS-element with adhesively bonded cap plates [6].CS-elements were immersed in alkaline pore water solu-tions of different temperatures during 18 months. It wasfound that the Arrhenius rate law could predict the elementstrength decrease. Due to the less harsh environment inpractice, the strength and stiffness decrease was found tobe acceptable, thereby making it possible to guaranteestructural safety and serviceability of the GFRP elementsafter 70 years of service.

3. Experimental investigation

3.1. Experimental beam description

Four beams with two-rib and four beams with three-ribTS-elements were poured. Fig. 5 shows one of the experi-mental beams with a two-rib TS-element. The arrangementof the GFRP elements and concrete reinforcement can beseen in Figs. 5 and 6. For both two- and three-rib beamsthe concrete cross-section was b = 400 mm wide andH = 240 mm deep. The total beam length of 1800 mm con-

Table 2Overview of experimental beams and measured results

Support type Loading mode No. of ribs Specimen Concretefcm (MPa

End (E) Moment (M) 2 M2E1 34.4M2E2 34.4

3 M3E1 32.7M3E2 32.7

Shear (S) 3 S3E1 32.7S3E2 32.7

Cantilever (C) Shear (S) 2 S2C1 34.4S2C2 34.4

sisted of 3 parts: the left (building interior) supported1000 mm long part, the 100 mm insulating joint, and theright (building exterior) 700 mm long cantilever part, whichwas loaded. The geometry was identical to the beams withhybrid-GFRP/steel joints [2], with two exceptions. First,the beam width had to be increased from 270 mm to400 mm due to the required stirrup diameter of 12 mm.This diameter resulted from an estimation of the transversetensile forces in the concrete due to the horizontal spread-ing of the concentrated loads introduced by the upper ribsinto the concrete. Second, only beams with 240 mm indepth were examined; 200 mm deep beams were notconsidered.

Also shown in Figs. 5 and 6 are the horizontal tubes inthe right part of the beam, which contained all wires fromthe strain gages (see next section). The concrete steel rein-forcement was of normal steel quality B500 according toSwiss Code SIA 262 [7]. Two times four beams were pouredsimultaneously; three 150 mm cubes were also cast to deter-mine the concrete cube strength after 28 days. The averageconcrete strengths were transformed into average prismstrengths, fcm, according to Ref. [7]. The resulting valuesare shown in Table 2.

4. Set-up, instrumentation and experimental program

The experimental set-up was the same as for the beamswith hybrid-GFRP/steel joints. The beams were fixedagainst uplift at the left end, as schematically shown inFig. 7. The right end of the left beam part was supportedon 100 mm wide steel plates in two different ways, bothof which can occur in facade constructions. In the arrange-ment designated ‘‘end support’’, the steel plate was alignedwith the beam edge thus providing a tri-axial stress state inthe concrete (confinement effect) at the section of loadintroduction from the lower GFRP flange of the CS-ele-ment. In the ‘‘cantilever support’’ arrangement the steel

strength)

Failure load measuredFu (kN)

Failure mode (first failure)

51.5 Anchorage failure supported side46.5 Anchorage failure loaded side65.9 Flange buckling CS-element62.1 Anchorage failure loaded side71.7 Shear failure in TS-element80.5 Shear failure in TS-element49.1 Anchorage failure loaded side55.1 Shear failure in TS-element

Fig. 7. Experimental set-up with different loading modes and supportconditions.

Fig. 8. Strain gage and wire arrangement in an upper 80 mm rib.


plate was shifted back by 100 mm thus preventing the addi-tional concrete strength due to confinement. The right can-tilever part of the beam was loaded by a hydraulic jackwith 200 kN capacity. Two different loading positions wereused, as shown in Fig. 7. At loading position e = 660 mmthe behavior of the beams was dominated by bendingmoment transfer, while at loading position e = 420 mmthe behavior was dominated by shear transfer (similar toRef. [2]). Table 2 gives an overview of all combinationsof the three parameters used: loading position (‘‘M’’ formoment and ‘‘S’’ for shear mode), support condition(‘‘E’’ for end and ‘‘C’’ for cantilever support), and numberof ribs (2 or 3). According to these letters the specimenswere denominated: M2E1, for example, designates beamnumber 1 with 2 ribs, end support, loaded in momentmode. As can be seen in Table 2, each combination wasexamined twice.

The beams tested in moment mode were only end sup-ported, because results from the beams with hybrid-GFRP/steel joints showed that similar behavior wasobtained with a cantilever support [2]. Since the upper ten-sile forces are a maximum in moment mode, the influenceof the number of ribs was studied in this mode. Anchoragefailure was expected for the M2E beams and concrete fail-ure at the lower edge of the loaded side for M3E beams(similar to the M240E beams in Ref. [2]). In shear mode,parameter combinations were chosen to give beams S3Eand S2C, where S3E corresponded to S240E in Ref. [2]and exhibited the highest shear load. For S3E a shear fail-ure in the TS-element was expected; for the S2C combina-tion, a concrete failure on the supported side at the loweredge was expected (similar to S240C in Ref. [2]). Calcula-tions showed that a tensile failure in the TS-element wouldhave never occurred due to the high element strength.

The vertical displacements at points A, B and C in Fig. 7were measured at each upper corner (two measurementlocations per point). Horizontal displacements were mea-sured at points B and D at each corner. In total, 10 dis-placement transducers with an accuracy of ±0.01 mmwere used. The CS-elements were instrumented as theywere for the hybrid joints, with twelve strain gages andtwo rosette gages each (see Ref. [2]). This layout was cho-sen to obtain the flow of forces in the element from the can-tilever to the supported side. The rosettes on the outer websof the CS-element consisted of three strain gages arrangedat 0� (horizontal axis), 45� and 90� to determine indirectly

the angle and magnitude of the principal stress and, there-fore, the portion of the shear force transferred by webs.

The two-rib TS-elements were instrumented with 61strain gages, while the three-rib elements were instru-mented with 81 gages to map the flow of forces from theloaded to the supported concrete part of the beam. Thestrain gage arrangements shown in Figs. 4 and 8 particu-larly enabled measurement of the strain decrease in theribs. Furthermore, two rosette gages were applied to theouter webs of the TS-element, as was done for the CS-ele-ment, to determine indirectly the portion of shear transferin the webs. The gages in the concrete part of the beamwere confined to the adhesive connections and the wireswere fed into the element cells through small holes in theflanges, as shown in Fig. 8. Inside the cells, the wires werearranged into two bundles protected by plastic tubing thatpenetrated the formwork and brought the wires to the exte-rior of the beam, as shown in Fig. 6. With this arrange-ment, all strain gages and wires were protected frommoisture and mechanical damage during pouring of theconcrete. All strain gages were delivered by HottingerBaldwin Messtechnik GmbH; the strain gages were of type6/120RY31 and the rosettes of type 6/120LY11. The loadwas measured with a load cell located between the hydrau-lic jack and the beams.

All beams were tested 28 days after pouring of the con-crete. The load was applied in four cycles ((1) 0–20 kN, (2)0–40 kN, (3) 0–60 kN and (4) 0–90 kN or ultimate failure)under displacement control at a rate of 5 mm/min. At eachcycle the crack pattern in the concrete was recorded. Theremainder of the paper concentrates on the results of thethird and fourth loading cycles.

5. Experimental results

5.1. Crack formation and failure modes

The first concrete cracks in the two-rib beams alloccurred on the upper side of the supported beam part dur-ing the first load cycle (0–20 kN, at 12 kN on average),while the first cracks in the three-rib beams also occurredin the supported part, but during the second load cycle


(0–40 kN, at 24 kN on average). Fig. 9 shows a typical fail-ure pattern (beam M2E1) with marked crack depths ateach load cycle (cycles up to 20, 40, 60 kN, R denotes thefailure cycle). The position of the TS-element is alsomarked. It can be seen that the first cracks appeared behindthe element end on the supported side, indicating a goodload transfer from the element to the parallel steel reinforc-ing bars. The beams tested in the moment mode developedfive cracks on the supported side, while beams tested in theshear mode developed three cracks.

The ultimate failure loads and failure modes of allbeams are given in Table 2. The two-rib beams showed sim-ilar failure loads (51 kN on average), as well as the lowestfailure loads. The three-rib beams tested in shear modeachieved the highest failure loads (76 kN on average), whilethe three-rib beams tested in the moment mode displayedmiddle failure load values (64 kN on average). In thetwo-rib beams, with exception of S2C2, anchorage failureoccurred on the loaded or supported side, as expected.That is, the TS-elements were pulled out of the concrete,as shown in Figs. 9 and 10. The steel stirrups betweenthe ribs showed high plastic deformations, see Fig. 10.The adhesive connection between the ribs and the TS-ele-ment never failed. Beam S2C2 showed a horizontal shearfailure in the webs of the TS-element in the joint gap. Vary-

Fig. 10. Anchorage failure of beam M2E1 on supported side, deformedsteel stirrups.

Fig. 9. Crack pattern of beam M2E1 (loading cycles and position ofTS-element marked, beam after demounting, supported side at right).

ing failure modes were seen in the three-rib beams: inM3E1 the upper flange of the CS-element failed first(upwards buckling), while M3E2 showed an anchoragefailure on the loaded side similar to the two-rib beams.Beams S3E1/2 failed due to a horizontal shear failure inthe webs of the TS-elements in the joint gaps, see Fig. 11;the other failures in the CS-element visible in Fig. 11 aresecondary failures.

6. Load–displacement response

Fig. 12 compares the load–displacement responses ofbeams M3E1 and S3E2 during the failure cycle. The beamin moment mode showed a considerably smaller ultimatefailure load than the beam in shear mode, but the deflection

Fig. 11. Parallel offset in joint and shear failure in TS-element (secondaryfailure of upper flange in CS-element).

Fig. 12. Measured load–displacement responses of beams M3E1 andS3E2.


developed was much bigger due to the more significantelongation of the TS-element and the four plasticallydeformed upper steel bars in the concrete. Comparing thecurves of points B and C of each loading mode indicatesthat the total cantilever beam deformation basically con-

Fig. 13. Measured axial strains on upper and l

Table 3Calculated shear forces transferred in upper TS and lower CS-element

Specimen Failure loadmeasuredFu (kN)

TS angle ofprincipalstress uTS (�)

CS angle ofprincipalstress uCS (�)

M2E1 51.5 45 35M2E2 46.5 45 46M3E1 65.9 50 30M3E2 62.1 42 35S3E1 71.7 44 20S3E2 80.5 46 42S2C1 49.1 45 26S2C2 55.1 46 23Average ± SD 60.3 ± 11.9 45 ± 2 32 ± 9

sisted of two different motions: a vertical offset in the jointand a rotation about the support point. The main deforma-tion in the joint was an almost parallel offset of the twoadjacent beam parts due to predominant shear deforma-tion, as visible in Figs. 9 and 11, leading to a rhomboidaldeformation of the CS- and TS-elements. The rotation inthe joint was comparatively small, particularly in the shearmode.

6.1. Load-strain response of compression/shear elements

Fig. 13 shows results of strain measurements on the CS-element of beam M2E2 during the failure cycle. This resultis representative for all other beams, independent of loadingmode or support conditions (exact locations of numberedgages are shown in Ref. [2]). The strain responses show thatthe upper flange exhibited high strains on the loaded side,which decreased towards the supported side. The lowerflange showed the opposite: low strains on the loaded sideand high strains on the supported side. Table 3, furthermore,gives the angles of principal stress, uCS, measured with therosette gages on the CS-elements (average values from bothsides). The angles with respect to the horizontal axis variedbetween 20� and 46�. All angles varied around the angle of

ower flange of CS-element in beam M2E2.

Shear TS/CSelement(%)

Shear TS elementVTS (kN)(sav (MPa))

Shear CS elementVCS (kN)(sav (MPa))

59/41 30.3 (25) 21.2 (18)49/51 22.8 (19) 23.7 (20)67/33 44.4 (37) 21.5 (18)56/44 34.9 (29) 27.2 (23)73/27 52.1 (43) 19.6 (16)53/47 43.1 (36) 37.4 (31)67/33 33.0 (28) 16.1 (13)71/29 39.1 (33) 16.0 (13)62/38 37.5 (31) 22.8 (19)


the geometric diagonal of the element (34�) and remainedalmost constant with increasing load up to failure.

6.2. Load-strain response of tensile/shear elements

Fig. 14 shows the axial strain distribution along theupper tensile flange of the TS-elements in beam M3E1and S3E1 at ultimate failure. Strain measurements beneaththe ribs and in the middle of the joint show how the strainsdecreased from the joint to the element ends, that is, howaxial stresses were transferred from the TS-element throughthe ribs to the upper steel reinforcement. Furthermore,Fig. 15 shows the corresponding distributions measuredon the lateral element sides of beam M3E1 (average fromboth sides) through the depth of the cross-sections at differ-ent locations (cross-sections at one end of each rib and inthe middle of the joint gap). The resulting distributionsare discussed in the next section together with the resultsof modeling. The distributions are representative for all

Fig. 14. Measured and calculated axial strain distribution along upper tensilefailure.

Fig. 15. Measured axial strain distribution through cross-section depth of

other beams, which showed similar results with similaraccuracy. Table 3 also gives the measured angles of princi-pal stress, uTS, measured with the rosette gages on theTS-elements (average values from both sides). The anglesof principal stress varied between 42� and 50�, independentof the beam type, and were bigger than the angle of the geo-metric diagonal (34�). Again, all angles remained almostconstant with increasing load up to failure.

7. Discussion

7.1. Transfer of shear forces in the joint

The shear forces, V, on the loaded side (V = Fu at ulti-mate failure) can be divided in two parts, VCS and VTS,transferred by the webs of the CS- and TS-elements tothe supported side according to Eq. (1):

V ¼ V TS þ V CS ¼ F u ð1Þ

flange of TS-element in beam M3E1 (left) and S3E1 (right), at ultimate

TS-element (average across width) in beam S3E1, at ultimate failure.


Since the joints showed an almost parallel offset of theadjacent beam parts, the division of shear forces can beestimated with the measured angles of principal stress inthe lower CS-webs, uCS, and upper TS-webs, uTS, asfollows: the vertical components of the principal stressescorrespond to the transferred shear forces, while the hori-zontal components of the principal stresses must be identi-cal in the lower CS and upper TS-webs assuming a paralleloffset and identical cross-sections:

V TS

tan uTS

¼ V CS

tan uCS

ðidentical horizontal componentsÞ ð2Þ

Combining Eqs. (1) and (2), the division of shear forces canbe calculated as follows:

V CS ¼ F u �tan uCS

tan uCS þ tan uTS

ð3Þ

V TS ¼ F u � V CS ð4Þ

The resulting shear forces at ultimate failure in the CS- andTS-elements are given in Table 3. The results show that62% (on average) of the shear forces are transferred inthe upper TS-element webs and 38% in the lower CS-ele-ment webs. Apparently, the TS-elements attracts a higherportion of the shear force due to the fixed support condi-tion on both sides, in contrast to the more flexible contactcap-plate connection of the CS-elements.

Table 3 also shows calculated average shear stresses, sav,in the element webs. The shear forces were divided by thearea of the four webs (4 · 5 · 60 = 1200 mm2). The designmanual provided by the pultruder gives a shear strength of25 MPa for similar shapes [5]. This value was exceeded inalmost all webs of the TS-elements, while only one beamshowed a higher value in the CS-webs (S3E2). Togetherwith beam M3E1, the S3E1/2 beams exhibited the highestshear stresses in the webs. For the TS-elements, whichdeveloped a shear failure (S3E1/2, S2C2, see Table 2), theshear stresses that resulted were up to 72% higher thanthe design manual shear strength (43, 36, 33 MPa) andtherefore led to the observed failure mode.

7.2. Transfer of bending moments in the joint

The bending moment at ultimate failure at the sup-ported joint edge is equal to Fu · e (e = lever arm of theload). Assuming that this moment is mainly transferred

Table 4Tensile forces and comparison of calculated and measured axial strains in the

Specimen T (kN) T 0 (kN) Ts (kN)

M2E1 198 30 228M2E2 178 23 201M3E1 253 44 297M3E2 238 35 273S3E1 175 52 227S3E2 197 43 239S2C1 120 33 153S2C2 135 39 173

in the 16 mm flanges of the upper TS-elements (in tension)and the lower CS-elements (in compression), the tensileforces in the upper 16 mm flange of the TS-element canbe estimated as follows:

T ¼ F u � eh

ð5Þ

h ¼ H � dT � dC ð6Þ

with h = 172 mm = inner lever arm between centers ofgravity of upper and lower 16 mm flange, H = 240 mm =section depth, dT = 35 + 12 + 8 = 55 mm = distance be-tween upper edge and center of gravity of upper 16 mmflange, and dC = 26/2 = 13 mm = distance between loweredge and center of gravity of equivalent contact area Ac0

(see Ref. [2]). The resulting tensile forces, T, in the upper16 mm flanges are listed in Table 4. The beams in momentmode showed approximately 38% higher tensile forces thanthe beams in shear mode, and the three-rib beams showedapproximately 37% higher tensile forces than the two-ribbeams.

The shear transfer of VTS in the upper TS-elementcaused a supplementary tensile force, T 0, in the upper16 mm flange at the supported joint edge. Due to the fixedend supports, a negative bending moment �T 0 Æ h 0 resultedin the TS-element at the supported joint edge and a positivemoment +T 0 Æ h 0 at the loaded joint edge. The distributionof this moment along the joint is linear from the negativemaximum at the supported side to the positive maximumat the loaded side with zero moment in the middle of thejoint. Assuming that these moments are mainly taken bythe upper and lower flanges of the TS-elements, T 0 canbe estimated as follows:

T 0 ¼ V TS � a2h0

ð7Þ

with: a = 100 mm = joint length, h 0 = 60–8–1.5 =50.5 mm = inner lever arm between centers of gravity ofupper 16 mm and lower 3 mm TS-flanges. The total tensileforce in the upper TS-flange, Ts, at the supported jointedge, therefore, is:

T s ¼ T þ T 0 ð8ÞAt the loaded joint edge, the tensile force, T, in the

upper TS-flange is reduced due to the positive moment tothe total tensile force, Tl, as follows:

upper TS-flange

Tl (kN) e(T) (%) e(Ts) (%) e(Tl) (%)

168 0.31 0.36 0.26156 0.28 0.31 0.24209 0.40 0.46 0.33204 0.37 0.43 0.32124 0.27 0.35 0.19154 0.31 0.37 0.24

87 0.19 0.24 0.1496 0.21 0.27 0.15

Table 5Concrete compression forces and stresses, and comparison of concreteconfinement factors from experiments and according to Swisscode SIA262

Specimen C (kN) rc (MPa) kc (–) kc,SIA (–)

M2E1 �198 �68 1.97 3.28M2E2 �178 �61 1.78 3.28M3E1 �253 �87 2.66 3.28M3E2 �238 �82 2.50 3.28S3E1 �175 �60 1.84 3.28S3E2 �197 �68 2.06 3.28S2C1 �120 �41 1.20 1.89S2C2 �135 �46 1.34 1.89


T l ¼ T � T 0 ð9Þ

The resulting tensile forces T 0, Ts and Tl are listed in Ta-ble 4. The supplementary tensile forces T 0 in the upperflange at the supported joint edge due to local TS-elementbending are approximately 20% of the tensile forces T fromthe global bending moment transfer.

From the total tensile forces in the upper TS-flanges, thecorresponding axial strains were calculated and comparedto the measured axial strains. The total axial tensile strainsat the supported joint edge, e(Ts), in the middle of the joint,e(T), and at the loaded joint edge, e(Tl), were calculated asfollows:

eðT sÞ ¼T s

AT � ET

ð10Þ

eðT Þ ¼ TAT � ET

ð11Þ

eðT lÞ ¼T l

AT � ET

ð12Þ

with: AT = 1600 mm2 = area of 16 mm tensile flange, andET = 40.5 GPa = tensile elastic modulus (see Table 1).The resulting axial strains at ultimate failure are listed inTable 4 and compared to the measured values in Fig. 14for beams M3E1 and S3E1. Considering the simplicity ofthe applied analytical model as well as the inherent scatterof single strain measurements, measured and calculated ax-ial tensile strains compare well over the joint length and al-low for an understanding of the combined tensile-shearload transfer mechanisms in the TS-element.

The compression force, C, in the lower 16 mm CS-flangedue to the global bending moment transfer is equal to thetensile force, T, in the upper 16 mm TS-flange:

C ¼ T ð13ÞFrom these compression forces, the compressive stresses inthe concrete, rc, were calculated according to Eq. (14):

rc ¼Cc

Ac0

ð14Þ

with: Ac0 = 112 · 26 = 2912 mm2 = equivalent contactarea, see Ref. [2]. Analogous to Ref. [2] for hybrid-GFRP/steel joints, a confinement factor, kc, can be calcu-lated according to Eq. (15)

kc ¼fcm

rc

ð15Þ

with fcm = uni-axial concrete prism strengths, given inTable 2. Table 5 shows the resulting compression stressesin the concrete, rc, and the resulting confinement factors,kc, which are compared to corresponding maximum con-finement factors, kc,SIA, calculated according to Swiss CodeSIA 262 [7] for the existent geometric configuration (seeRef. [2]). The comparison shows that the maximum possi-ble confinement factors were not reached and thereforeconcrete failure never occurred, as observed in the experi-ments. This result was different from the results of the hy-

brid-GFRP beams, where concrete failure always occurredin the moment mode and shear mode with cantilever sup-port (see [2]).

7.3. Anchorage of tensile/shear element

All two-rib elements were pulled out of the concrete, asexpected, with exception of that of beam S2C2 where shearfailure in the element occurred first (see Table 2). For thethree-rib elements, only the element in beam M3E2 waspulled out. Post-experimental inspection, however, showeda misalignment of the first steel stirrups in beam M3E2,which prevented the retention of the element. It can beassumed that without this misalignment beam M3E2would have shown a similar failure mode as beam M3E1,which did not exhibit anchorage failure.

The anchorage on the supported side was always moreloaded – with Ts according to Eq. (8) – than the anchorageon the loaded side (loaded with Tl, see Eq. (9)). Fig. 14 alsoshows the strain development in the two anchorage zonesof the TS-element in the three-rib beams M3E1 and S3E1at ultimate failure. The calculated curves are based onthe assumption that each rib anchors the same amount offorce, that is, 33% (in three-rib beams) of Ts or Tl. Insidethe bonded ribs the load transfer was approximated as lin-ear, even though it was known that a stress peak occurs atthe load transfer section from the rib to the concrete. Theresulting curves from modeling and measurements com-pare well considering the scatter that can occur in singlestrain measurements. The assumption of an almost linear,step-wise load transfer from the TS-element to the parallelupper steel reinforcement bars was confirmed by theseresults. The agreement of calculated and measured axialstrains also confirms that the simple analytical model usedis accurate enough to describe the complex load-carryingbehavior of the beams.

Fig. 15 shows the axial tensile stress distributionsthrough the thickness of the cross-sections. The threecross-sections shown in the middle (at the supported jointedge (left), at the middle of the joint and at the loaded jointedge (right)) show the influence of the local bending due toa shear transfer. The strain distribution in the middle cross-section is not influenced by the local bending moment,

Table 6Comparison of results from hybrid GFRP/steel and all-GFRP joints

Specimen Joint type Failure load measured Fu (kN) Shear portion CS-element (%) Angle of principal stress uCS (�)

M3E1 All-GFRP 65.9 33 30M3E2 All-GFRP 62.1 44 35M240E1 Hybrid-GFRP 54.4 0 10M240E2 Hybrid-GFRP 64.6 10 8

S3E1 All-GFRP 71.7 27 20S3E2 All-GFRP 80.5 47 42S240E3 Hybrid-GFRP 81.6 57 38


while in the adjacent cross-sections, the strains from thenegative and positive moments are superposed.

7.4. Comparison of all-FRP and hybrid-FRP joint

performance

Table 6 compares the three-rib beams with the corre-sponding hybrid GFRP/steel joint beams described in [2].The shear and upper tensile steel bars in the hybrid-GFRPjoint were replaced by the TS-element in the all-GFRPjoint. The ultimate failure loads and system stiffnesses ofboth systems compared well in both loading modes. Sheartransfer in shear mode was similar with similar portions oftotal shear and angles of principal stress in the CS-element.In moment mode, however, the CS-element in the hybridjoint contributed only negligibly to the shear load transferin contrast to the all-GFRP joint, where no differences ofparticipation was noticed between the two loading modes.

Both systems showed similar ductile behavior, eventhough the all-GFRP joint consists of purely brittle mate-rials. Fig. 16 shows the load–displacement response ofbeam M240E2 (hybrid-GFRP joint) and correspondingbeam M3E2 (all-GFRP joint) during cyclic loading. Thetwo curves compare well. Beams with hybrid joints showedclassic ductile concrete failure during yielding of the steelbars in the joint [2]. In beams with all-GFRP joints, ductil-ity was provided by three different mechanisms: (a) yielding

Fig. 16. Comparison of load–displacement response of all-GFRP jointbeam M3E1 and hybrid-GFRP/steel joint beam M240E2.

of the upper steel bars in the concrete on the supported sideof the joint (beams M2E1/2, M3E1/2, S3E), (b) plasticdeformation of the steel stirrups during anchorage failure(beams M2E1/2, M3E2, S2C1), and (c) pseudo-ductileGFRP failure in the CS-element of beam M3E1 andTS-element of beams S3E1/2 and S2C2 (pseudo-ductilitydefined according to Ref. [8]). However, as it can be seenin the third cycle, a certain visco-elastic behavior of theGFRP components occurred, that is, the residual deforma-tions after the third cycle are partially recovered.

8. Conclusions

An existing hybrid-GFRP/steel joint for load transferand thermal insulation in concrete slab structures was fur-ther developed to an all-GFRP joint by replacing theremaining steel bars with a pultruded GFRP tensile/shearelement anchored in both sides of the joint in the concrete.The GFRP compression/shear element is the same in bothjoints. The quasi-static performance of the new joint wasinvestigated through full-scale experiments on cantileverbeam structures and analytical modeling. The followingconclusions can be drawn from this work:

(1) The new all-GFRP joint with a three-rib tensile/shearelement provides similar ductile performance to thehybrid-GFRP/steel joint. Failure, however, occursin the GFRP elements in contrast to the hybrid jointswhere concrete failure always occurred during yield-ing of the steel bars.

(2) The transfer of bending moments occurs throughcompression forces in the lower flange of the com-pression/shear element and tensile forces in the upperflange of the tensile/shear element, similar to thehybrid joints that transfer tensile forces in the steelbars.

(3) The anchorage of the tensile/shear element in theconcrete with two or three ribs adhesively bondedto the element on each joint side provides a linear,step-wise transfer of the tensile forces from the ele-ment to the parallel steel reinforcement bars in theconcrete.

(4) The shear force transfer is shared between bothGFRP elements. The tensile/shear element attracts50% more shear force due to the fixed end supports


and transfer occurs through local bending, whichincreases the tensile forces in the upper flange by20% at the supported joint edge.

(5) Simple analytical models allow the joint behavior tobe described and compare well with measurements.

Compared to the hybrid-GFRP/steel joint, the new all-GFRP joint allows for a further reduction of approxi-mately 50% of the linear thermal bridge allowance. Furtherstudies are ongoing with regard to the elaboration of adetailed model of the joint behavior (including concretecracking and joint rotation) as a basis for a design methodfor application in practice.

Acknowledgements

The authors would like to thank the Swiss InnovationPromotion Agency (Contract Number 6278.1KTS) andSFS Locher AG (Switzerland) for supporting this re-search.

References

[1] Schweizerische Normen-Vereinigung, Swiss Code SIA 380/1: Thermi-sche Energie im Hochbau. Schweizerischer Ingenieur- und Architekt-enverein (SIA), Zurich, Switzerland, 2001.

[2] Keller T, Riebel F, Zhou A. Multifunctional hybrid GFRP/steel jointfor concrete slab structures. J Compos Construct, 2006.

[3] SFS Locher: SFS Locher. Information brochure/price list – isolan andisolan plus, Register 8. Available from http://www.sfslocher.biz/web/sfsloh_d.nsf/vwFraSet/bausysteme (3 June, 2005).

[4] Keller T, Vallee T. Adhesively bonded lap joints from pultrudedGFRP profiles, Part II: Joint strength prediction. Composites Part B2005;36/4:341–50.

[5] Fiberline Composites A/S. The Fiberline design manual, 2nd ed.Kolding, Denmark. Available from: http://www.fiberline.com (3 June,2005).

[6] Riebel F, Keller T. Long-term compression performance of apultruded GFRP element exposed to concrete pore water solution. JCompos Construct, submitted for publication.

[7] Schweizerische Normen-Vereinigung. Swiss Code SIA 262: Betonbau.Schweizerischer Ingenieur- und Architektenverein (SIA), Zurich,Switzerland, 2003.

[8] Keller T, De Castro J. System ductility and redundancy of FRP beamstructures with ductile adhesive joints. Composites Part B 2005;36/8:586–96.

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multifunctional all-gfrp joint for concrete slab structures

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