design, manufacture and testing of a new hybrid column

Construction and Building Materials, Vol. 12, No. 1, pp. 39]49, 1998Q 1998 Elsevier Science Ltd

Printed in Great Britain. All rights reserved0950]0618r98 $19.00 q 0.00

( )PII:S0950–0618 97 00073–1

Design, manufacture and testing of a newhybrid column

Amir MirmiranU†, Michel SamaanU, Suramy CabreraU and Mohsen Shahawy‡

UDepartment of Civil and Environmental Engineering, University of Central Florida,POBox 162450, Orlando, FL 32816-2450, USA‡Structural Research Center, Florida Department of Transportation, Tallahassee,FL 32310, USA

Received 15 December 1996; revised 1 September 1997; accepted 2 November1997

( )Hybrid structures with concrete and fiber reinforced plastics FRP are regarded as efficient systemswith pseudo-ductile characteristics and high stiffness and strength. This article presents the design,fabrication and testing of a new hybrid column made of concrete-filled FRP tube. The tube acts asformwork, protective jacket, confinement and shear and flexural reinforcement. Longitudinal andtransverse stiffening ribs in the tube act as shear connectors and load distribution mechanism.Excellent shear performance was achieved in beam shear tests. Moreover, hybrid columns incompression control proved as strong in moment capacity as equivalent RC columns with over 5%rebars. Q 1998 Elsevier Science Ltd. All rights reserved.

Keywords: column; concrete; confinement

Introduction

A hybrid column is proposed that consists of a fiberŽ .reinforced plastic FRP tube filled with concrete, for

either cast-in-place or precast construction. The pri-mary objective is to utilize full section enclosure andconfinement of concrete by FRP with or without inter-nal reinforcement to enhance load carrying capacityand durability of the column. The benefits are twofold;protection of concrete against harsh environmental ef-fects and increasing the column’s strength by confiningthe concrete core. The concept of confining concretewith FRP materials has been used for strengtheningand retrofitting in earthquake prone zones since themid-1980s and often implies wrapping and bondingbi-directional fiber straps or belts1,2, thick FRP cables3

or precured composite shells3,4 around existingcolumns. The columns retrofitted as such have gainedsubstantial enhancements in their strength, ductilityand shear performance and have withstood the 1994Northridge earthquake with no sustained damage4.Previous studies have shown that lateral confinement

†Tel.: q1 407 8232841; fax: q1 407 8233315; e-mail:[email protected].

of concrete increases its ultimate compressive strainfrom a value of 0.005 in an unconfined section to avalue of 0.03 or higher 5. This significantly enhancesductility of the section. The radial confining stresses inconcrete are balanced by tensile hoop stresses devel-oped in the exterior shell, for which a filament-woundFRP cylindrical section is ideal. Additional cost savingsare realized by reducing the cross-sectional area re-quired for the same design load. Moreover, use of FRPforms to pour concrete and retaining the form as partof the structural system would save labor costs andconstruction time. The proposed system may provemost economical as column for high-rise buildings,parking garages and bridge piers, as well as precastpiles.

The Florida Department of Transportation has spon-sored a detailed study on the feasibility of hybridFRP-concrete members6. Experimental investigationinto the performance of hybrid columns included: uni-axial compression tests to establish the effect of con-finement with FRP on strength and ductility of con-crete; uniaxial cyclic tests to evaluate the unloadingand reloading response of hybrid columns; split-cylin-der tests to estimate tensile strength of the compositesection; split-disk tests to evaluate the strength and

39

A new hybrid column: A. Mirmiran et al.40

stiffness of FRP tubes; beam shear tests to determinethe contribution of FRP tube to the shear resistance ofhybrid columns and to investigate the effectiveness ofshear transfer mechanisms; beam]column tests to gen-erate an experimental interaction diagram for hybridcolumns; square column tests to evaluate the effect ofcross-sectional shape on confinement of concrete; andslenderness tests to estimate the length effect on load-carrying capacity of the column. In this article, first thedesign and manufacture of square FRP tubes withshear ribs are discussed and then the results of beamshear tests and beam]column tests are presented.

Design and manufacture of the mandrel

Similar to pressure vessels, FRP tubes for hybridcolumns can be easily fabricated by the filament-wind-ing process. In this process, fibers are impregnated asthey run through a resin bath and are then wound on arotating mold, the mandrel. The cross-section of thetube depends on that of the mandrel. There is usually adrift angle built into the mandrel for easy separation ofthe final product. These tubes, however, lack in anymechanical bond with the concrete core. Since withoutsuch bond composite action is less likely to develop, anew type of FRP tube was designed with bi-directionalshear connector ribs on the interior surface of the tube.It should be noted that the tube is intended as bothlongitudinal and transverse reinforcement for concrete.For the tube to act as longitudinal reinforcement, notonly its detailing at the column ends is important, butalso its bond with the concrete core is necessarythroughout the length of the column. For the tube toact mainly as transverse shear reinforcement, a sheartransfer mechanism is still necessary, otherwise thetruss analogy will not be applicable. The proposedshear connectors are normally used in flight and ma-rine structures as longitudinal and transverse stiffenersfor FRP shells. Figure 1 shows a fuselage component, inwhich the longitudinal stiffeners or stringers serve thefollowing purposes7: they resist bending and axial loads

Ž .along with the skin i.e. the shell ; they divide the skininto small panels, thereby increasing its buckling andcompressive strength; and finally, they help arrest thegrowth of cracks in the skin by accepting tensile loads

Figure 1 Longitudinal and transverse stiffeners in a fuselage com-ponent7

that the skin can no longer carry. The transverse ribsare, on the other hand, used to: maintain the cross-sec-tional shape of the skin; re-distribute stresses aroundstructural discontinuities; provide end restraints for thelongitudinal ribs; provide edge restraints for the skinpanels, thereby increasing their plate buckling stress;and finally, act with the skin in resisting the hoopstresses.

In order to make the shear stiffeners, the exteriorsurface of the mandrel needs to have the oppositeshape of the desired ribs. However, it is understoodthat unless such a mandrel is collapsible, it will not bepossible to remove the final product from the mandrelbecause of the interlocking connections and also due toshrinkage of the resin. Therefore, a collapsible mandrel

Ž .was designed and fabricated from four 4 aluminumangles of 2.5=2.5=3r8 inches each with a length of

Ž .5.5 feet 66 inches . The angles were assembled togetherwith a 1-inch gap in between to form the shape of a6=6-inch section. The assembly was held in place on a1-inch diameter, 12-foot long driving shaft with two endplates and three interior supporting plates in between.The shaft would be driven by a filament winding ma-chine. The entire mandrel is easily erected and disman-tled. On each face of the aluminum angles a series of2=1.8=0.25 inch maple wood plates were mountedwith 0.5 inches of spacing in between. Each wood plate

Ž .was secured in place with four 4 screws. The woodplates were properly waxed to avoid adhesion to theresin that was applied during the filament-windingprocess. Figure 2 shows the collapsible aluminum man-drel with the wood plates.

Design and fabrication of the specimens

For the shear ribs, a special putty made of polyesterresin and silica fume was mixed to a pre-specifiedviscosity. The mix was further strengthened by addingchopped glass fibers of approximately 1% by volume toincrease the shear strength of the ribs. After mixing the

Ž .putty with chopped fibers, 1.5% by volume methyl

Figure 2 Collapsible aluminum mandrel with the wood plates

A new hybrid column: A. Mirmiran et al. 41

Ž .ethyl ketone peroxide MEKP was added as catalyst.The 1-inch gaps between the aluminum angles werecovered with masking tapes. As shown in Figure 3, theputty was then applied to the spaces in between thewood plates to achieve a smooth and planar surface. Ittakes approximately 10]15 min before the putty be-comes too hard to work with. It then takes another30]45 min for the putty to fully cure. Curing of theputty is directly related to the ambient temperatureand the amount of catalyst. Addition of more catalystdecreases both curing time and working time. Thecuring process is often associated with some heat,change of color from pink to brown and considerableshrinkage. The shrinkage of the putty increases when ahigher percentage of catalyst is used, or when theambient temperature is raised. Excessive shrinkage maylead to the putty bulging out of the spaces in betweenthe wood plates, resulting in a larger cross-sectionalarea of the tube. After the putty was cured, a layer of

Ž .polyester resin with 1.5% by volume MEKP as cata-lyst was applied and a 24-oz bi-directional fiberglasswoven roving was laid on each of the four sides of themandrel to provide the axial reinforcement in the

Ž .structure of the tube see Figure 4 . The roving wasfurther soaked and saturated with the resin to removeall air voids from underneath and within the roving

Figure 3 Applying the putty to the spaces between wood plates

Figure 4 Placement of the fiberglass woven roving

Figure 5 Saturation and soaking of the fiberglass woven roving withresin

Ž .see Figure 5 . It should be noted that the tubes madefor the beam shear tests did not include any fabric,except for one of the tubes that included a choppedstrand mat, instead. Upon completion of this phase, theregular winding process was carried out on a nowpractically smooth and planar surface of a square man-drel. Figure 6 shows the mandrel on the filament-wind-ing machine. After the winding process was completed,the mandrel was placed horizontally on a revolvingsystem with hot air blowing from its bottom to speed upthe curing process which would normally take approxi-mately 30]45 min. The mandrel was then placed inroom temperature for approximately 12]15 h for thetube to completely cure. The next morning, the man-drel was collapsed inward to remove the fabricatedtube. Figure 7 shows the interior shear ribs of the FRPtubes. The longitudinal ribs are 1.65 inches wide andthe transverse ribs are 0.75 inches wide. All shear ribsare 0.25 inches thick.

Beam shear tests

In order to examine the effect of mechanical bond onthe shear transfer between the tube and the concretecore, two series of beam shear tests were conducted,one with no shear ribs and the other with the ribs. Allspecimens of both series were tested in a four-pointloading set-up with 18-inch span length. In this section,first a review of previous work and the analytical shearstrength of hybrid beams are discussed and then thetwo series of tests are presented.

Re¨iew of pre¨ious work

Use of steel or FRP jackets for shear repair of concretebeams and columns has become popular in recentyears. Chai8 addressed the flexural performance ofsteel jacketed columns and showed the effectiveness ofa steel jacket as a confining mechanism. Priestley etal.9,10 studied the shear enhancement due to steel jacketretrofitting of RC bridge columns and suggested that


Figure 6 Filament winding of the square tube

the jacket may conservatively be idealized as a series ofindependent closely spaced peripheral hoops withthickness and spacing equivalent to the jacket thick-ness. The experiments included a total of 14 large-scalecircular and rectangular columns. They proposed thefollowing relationship for the shear strength enhance-ment provided by a circular jacket

Ž . Ž .V s0.865 p f t D y t 1j j j j j

where f sstrength of the jacket, t s thickness of thej jjacket and D soutside diameter of the jacket. Injderiving the above relationship, a 308 inclination anglewas assumed for the diagonal struts. The angle wasverified by tests on steel jackets10 and compositejackets11. They further recommended using an ellipticaljacket for rectangular columns and derived the fol-lowing relations for the strong and weak directions ofthe elliptical jacket

Bp jŽ .V s3.46 f t D y t 1y 1yj j j j j ž /4 Dj

Strong direction

Dp jŽ .V s3.46 f t B y t 1y 1yj j j j j ž /4 Bj

Ž .Weak direction 2

Figure 7 Interior shear ribs of the FRP tubes

where B and D are the outside dimensions of thej jelliptical jacket in the short and long directions, respec-tively.

Al-Sulaimani et al.12 investigated the effectiveness ofbonded FRP plates on the shear capacity of RC beams.The experiments included 16 RC beams that weredeliberately designed to be deficient in shear capacity.All beams were initially loaded to a pre-determinedlevel. The damaged beams were then divided into fourgroups according to the form of repair. The four repair

Ž .categories were as follows: 1 control beams with noŽ .shear repair; 2 beams repaired by shear strips, i.e.

Ž .strips of FRP plates; 3 beams repaired by shear wings,Ž .i.e. FRP plates on the sides only; and 4 beams re-

paired by U-jackets. In each category, half of thebeams were only strengthened for shear, while theother half were further strengthened for flexure bybonding an FRP plate on the soffit of the beam. Thebeams were then loaded to failure. In general, thebeams repaired with the U-jackets showed the bestshear performance. This was attributed to the fact thatthe continuity rendered by the geometry of the jacketminimized the effect of stress concentrations. The shearstrength for U-jacketed beams was so high that eventhe beams with flexural FRP plates failed in flexure.

Chajes et al.13 conducted an experiment on the shearstrength of RC beams with externally applied compos-ite fabrics. A total of 12 T-beams were tested underfour-point loading. Of these, four were control speci-mens with no fiber-wraps, two were wrapped with 0]908aramid, two with 0]908 E-glass, two with 0]908 carbonand two with"458 carbon fabrics. While all beamsfailed in shear, the externally wrapped beams displayedan increase of 60]150% in load-carrying capacity overthe control specimens, with the beams wrapped with"458 carbon fabrics showing the highest increase instrength. They also used truss analogy and proposedthe following relationships for the shear strength of thefabric

V sE A de for 0y908 weavej j j ¨cu

' Ž .V sE A de 2 for "458 weave 3j j j ¨cu

where E smodulus of elasticity of resin-impregnatedjfabric, A sarea of fabric shear reinforcement per unitjlength of the beam, dsdepth of the tension reinforce-ment and e sultimate vertical tensile strain of con-¨cucrete.

Analytical shear strength of FRP-encased concrete

Because of the generally sudden and often catastrophicnature of shear failures, it is necessary to developmodels for adequate prediction of the shear strength ofhybrid columns. It is generally assumed that the exter-nal plate acts as a continuous array of shear stirrups.Figure 8 shows a cracked hybrid beam with rectangularcross-section. If the tube is made of"f angle pliesand if the angle of shear failure plane is assumed to beu , the shear resistance by the tube, V , is given byj


Figure 8 Shear crack in a rectangular FRP beam with"f angleplies

2 f t d cos fj j Ž .V s 4j tan u

where f sstrength of the plies in the direction ofjfibers, t s thickness of the tube and dsdepth of thejcross-section. The above equation is derived based on a‘netting analysis’14 in which the stresses in an FRPstructure are predicted by neglecting the contributionof the resin. The technique applies static equilibriumwith no consideration of strain compatibility14. Theapproach is believed to produce a conservative lowerbound value, since contribution of the resin is ne-glected.

The shear strength provided by a circular tube is,however, somewhat different. Chai8 addressed the shearstrength of circular retrofitted columns with steel jack-ets. He assumed that the jacket will yield in a state ofhoop stress with a 458 failure plane in concrete. Asimilar approach is adopted here with slight modifica-tion to account for different material and windingangle of the tube. Figure 9 shows the shear failureplane in a cylindrical FRP tube. For an infinitesimallength of the tube, d z, the shear force resisted by thetube is

Ž .dV s2 f t cos f sin a d z 5j j j

From the geometry, we can write

yr cos a Ž .zs 6tan u

from which

r sin a Ž .d zs da 7tan u

Ž . Ž .Substituting Eq. 7 into Eq. 5 and integrating for abetween 0 and p , the shear strength of the tube can befound as

Ž .p f t D y t cos fj j j j Ž .V s 8j 2 tan u

The total shear strength of the hybrid section canŽ .then be computed by adding the results from Eq. 4 or

Figure 9 Shear crack in a circular FRP beam with"f angle plies

Ž .Eq. 8 to the shear strength of concrete core, V , whichccan be estimated from the ACI Building Code15 as

X Ž .V s2 f bd 9'c c

where f X s28-day compressive strength of concretecand bswidth of the section. It should be noted that

Ž .Eq. 9 is generally quite conservative. This point wasconfirmed in the beam shear tests for the controlbeams in the present study as well as studies byothers9,10. Moreover, full enclosure of the section helpsincrease shear strength of the concrete core by merelycontaining the cracks. At the first onset of cracks inconcrete, the load would be transferred to the tube,which would retard any further increase in the width ofthe cracks if no slippage is allowed to occur. If, how-ever, slippage occurs, the tube will simply contain thecracks rather than working as shear stirrups. In thatevent, the above equations will not provide an accurateestimate of the shear strength.

Series 1: tubes without shear ribs

Series 1 consisted of eight 22-inch-long concrete beamswith a 6=6-inch cross-section in four groups as shownin Figure 10. For each group, two similar specimenswere prepared and tested. The average strength of theconcrete core was determined to be 4476 psi fromcontrol cylinders. All beams were reinforced with twoNo. 4 bars on the compression side and three No. 4bars on the tension side. Rebars were affixed to the


Figure 10 Specimen layout for Series 1

inside of the tubes with thin steel wires tightly tied. Noshear stirrup was provided in any of the specimens. The

Ž .control beams Group A were designed with a flexuralstrength of approximately nine times their shearstrength to ensure shear failure. The FRP tubes con-sisted of seven angle-plies of polyester resin with uni-directional E-glass fibers wound at"158. The fiberrovings were Vetrotex Certain Teed 67BrR099r450-Yield E-glass and the polyester resin was Dion FRr33-611r6692T manufactured by Ashland Chemical Com-pany. A 0.02-inch-thick layer of uni-directional weavecarbon fabric made of AMOCO THORNEL yarns wasglued to the inside face of the FRP tubes only forspecimens of Groups C and D to further increase theirflexural strength. Table 1 shows the manufacturer’s

Ždata for the resin-impregnated glass fibers ASTM D-.2343 , the polyester resin and the carbon sheet. Prior

to testing of the specimens of Group D, the top face ofthe tube was cut and removed to examine the effect ofcontinuity and full-section enclosure on shear response.

All beams were instrumented with rosettes at mid-height of the section along the 458 line between theload and the support on both sides. Also, strain gageswere placed on the top face of the beams at mid-span.Deflections were monitored by three LVDTs, one atthe mid-span and one under each loading point. Test-ing was carried out in a displacement control mode.

Throughout the test, significant popping noise could beheard which was attributed to cracking of concrete andits de-bonding from the tube. As the loading pro-gressed, white lines developed in the tube indicatingcracking and flow of the resin. It was also possible tosee the concrete slipping out of the tube and breakingoff at each end. Near the end of the loading process,vertical faces of the tubes began buckling outwardsunder the load. At the time of failure, an average of0.25]0.5 inches of concrete had slipped out of the tube

Ž .and in some specimens broke off at either end. Thisindicated very little bond between the concrete and thetube. Specimens of Group C were the only ones toclearly show flexural cracks. It is possible that theweave of the carbon sheet helped partially inhibit slip-page on the tension face of concrete resulting in amore successful transfer of load between the concreteand the tube. Generally, failure of the tube occurredat"158 which lined up with the fiber orientation in thetube. After removing the tube from the failed speci-mens, one could clearly see extensive shear cracks inconcrete. The tube did not function as shear stirrups inan RC beam. However, the shear strength was in-creased considerably due to crack containment. This isclear by comparing the failure load for control beamswith those of the jacketed beams as appear in Table 2.

ŽIt was evident from specimens of Group D with no top

Table 1 Mechanical properties of glass fibers, polyester resin and carbon sheet

Property 450-Yield E-glass Polyester resin Thornelcarbon fiber T-300 12k

Specific gravity 2.58 1.41 1.77Ž . Ž . Ž . Ž .Tensile strength, MPa ksi 2186 317 72 10.4 3650 530Ž . Ž . Ž . Ž .Tensile modulus, MPa ksi 69,640 10,100 4344 630 231,000 33,500

Ž . Ž . Ž . Ž .Shear modulus, MPa ksi 30,130 4370 1600 232 4970 700Poisson’s ratio 0.22 0.36 0.20


.plates that part of the enhancement in shear strengthseen for other groups comes form the ability of the

Ž .tube to contain rather than confine the concreteallowing it to withstand higher loads. For Group Dspecimens, the load was applied directly onto the con-crete and in fact the tube did not show any cracks,indicating that no load was transferred to the tube byconcrete. Table 2 shows the increase in the shearstrength for the jacketed specimens, as well as a com-parison with the theoretical shear values. Theoreticalshear strength of control specimens is calculated from

Ž .Eq. 9 as proposed by the ACI, while shear strength ofhybrid specimens is calculated by adding results of Eq.Ž . Ž . Ž .4 and Eq. 9 . In Eq. 4 , directional strength of FRPŽ . 6f cos f is assumed as 76 ksi, thickness of the tubejis 0.063 inches and depth of concrete section is as-sumed to be 6 inches. As shown from the table, theACI equation underestimates shear strength of con-

Ž .crete beams. On the other hand, Eq. 4 overestimatesshear contribution of the jacket mainly because of theexcessive slippage and lack of any shear transfer mech-anism. Although, crack containment resulted in con-siderable shear enhancement of concrete beams overthe control specimens, without a shear transfer mech-anism that would arrest any slippage full contribution

Ž .of the jacket as estimated by Eq. 4 cannot be realized.

Series 2: tubes with shear ribs

In order to correct the slippage problem encounteredwith the beams of Series 1, the FRP tubes with shearconnectors on all four interior faces were used. Asshown in Figure 11, Series 2 consisted of eight 24-inch-long beams with a 6-inch2 cross-section in four groups,where two specimens were tested for each group. Theaverage strength of concrete core was determined to be4722 psi from control cylinders. The control specimensof Group E had a glass sheet glued onto their tensionside as flexural reinforcement. No steel reinforcementwas provided in any of the specimens for either flexureor shear. Specimens of Group F were encased in a0.135-inch-thick FRP tube made of 15 plies. Specimensof Group G were similar to those of Group F, exceptthat a chopped strand mat was built into the tube priorto the winding process. Specimens of Group H wereencased in a 0.054-inch-thick FRP tube with six pliesand no mat. Instrumentation and test set-up was simi-lar to Series 1, except that additional strain gages wereattached to the bottom face of all beams at mid-span.The control specimens failed in flexure due to bondfailure between the glass sheet and the concrete. TheFRP-encased beams all failed in flexure, with speci-mens F1, H1 and H2 failing at mid-span and specimensF2, G1 and G2 failing under one of the loads. This wasattributed to the unevenness of the tube surface whichresulted in uneven distribution of the load, biaxialbending and some torsion. Table 3 shows the increasein the shear strength for the jacketed specimens, aswell as a comparison with the theoretical shear values.Theoretical shear strength of control specimens is cal-

Ž .culated from Eq. 9 , while shear strength of hybrid

Table 2 Test results for Series 1

Specimen Experimental Percent increase Theoretical P rPTheo ExpŽ . Ž .P kips over control beams P kipsmax max

A1 9.96 } 4.82 0.5A2 11.93 } 4.82 0.4B1 20.74 89.49 62.27 3.0B2 25.55 133.44 62.27 2.4C1 23.74 116.90 62.27 2.6C2 19.12 74.69 62.27 3.3D1 14.18 29.56 62.27 4.4D2 17.45 59.43 62.27 3.6

Ž .specimens is calculated by adding results of Eq. 4 andŽ . Ž . ŽEq. 9 . In Eq. 4 , directional strength of FRP f cosj

. 6 Ž .f is assumed as 76 ksi for six-layer tubes with ribsŽ .and 93 ksi for 15-layer tubes with ribs , thickness of

the tube is 0.054 and 0.135 inches for the six and15-plies, respectively and depth of concrete section isassumed to be 6 inches. In these calculations, effect ofribs is not quantified separately. No slippage was evi-dent in any of the beams. As a result, flexural crackingwas clearly visible in most of the specimens. Uponremoval of the tubes, there was no evidence of exten-sive shear cracks present in concrete, unlike specimensof Series 1. The shear connectors had effectively re-tarded slippage and allowed for the concrete and thetube to act together as a composite system. Since allspecimens failed in flexure, theoretical values of shearstrength in Table 3 were much higher than the failureloads. Figures 12 and 13 show graphs of the top andbottom strains for all specimens. It is clear that bothstrength and ductility of concrete beams increase as aresult of crack containment in the hybrid section. It isalso of interest to note that the hybrid specimensexhibited the same response as the control specimens

Žup to the point where the plain concrete fails i.e..rupture under flexural tension . However, after this

point, a clear change in the slope occurs. All hybridspecimens generally show the same second slope intheir load]strain response curves. Figure 14 shows themoment]curvature graphs for all specimens of Series 2at their mid-span. The curvatures are calculated basedon the top and bottom strains. Hybrid specimensdemonstrate considerable ductility without any steelreinforcement.

In comparison, although shear demands for speci-mens of Series 1 and 2 were quite different, it was clearthat presence of shear ribs could arrest any slippageand allow full utilization of the tube as a continuousarray of shear stirrups. For Series 1, since slippageoccurred, the tube did not act as shear stirrup and,therefore, theoretical values of shear strength were notreached. On the other hand, for specimens of Series 2,shear capacity was raised so high that all specimensfailed in flexure.

Beam–column testsIn addition to the short beam shear tests, a total of fivelong beam]column tests were performed. The tests


Table 3 Test results for Series 2

Specimen Experimental Percent increase Theoretical P P rPmax Theo ExpŽ . Ž .P kips over control beams kipsmax

E1 4.54 } 4.95 1.1E2 4.11 } 4.95 1.2F1 27.55 537 155.61 5.6F2 19.85 359 155.61 7.8G1 10.64 146 155.61 14.6G2 10.80 150 155.61 14.4H1 24.68 471 54.20 6.3H2 24.32 462 54.20 6.4

Figure 11 Specimen layout for Series 2

Figure 12 Load vs. top strain for specimens of Series 2


Figure 13 Load vs. bottom strain for specimens of Series 2

were devised to develop an experimental interactiondiagram for the hybrid columns. Three types of tests

Žwere conducted; pure axial load with no flexure one. Žspecimen , pure flexure with no axial load one speci-

. Žmen and combined axial]flexural tests three speci-.mens . All specimens were 7=7=52 inches, with a

span length of 48 inches between the supports. TheFRP tube consisted of shear connectors, a 24-oz bi-di-rectional fiberglass woven roving on all four sides and15 angle-plies of"158 winding. The average strengthof concrete core was determined to be 2720 psi fromcontrol cylinders. All specimens were tested in a hori-zontal position between two reaction walls, a strongfloor and a 36-inch steel beam at the top. Figure 15shows the schematic of the test setup for combinedaxial and flexural loads. The axial]flexural tests wereconducted at three different levels of axial loads, whichcorresponded to one-eighth, one-half and three-quarters of the maximum axial load. First, the speci-men was loaded axially to the desired level. Then, whilemaintaining the axial load level, the specimen wassubjected to a monotonically increasing transverse loadsimilar to a four-point flexure test. For the first level ofaxial load, a tension control failure took place, whilethe other two tests resulted in a compression controlfailure. Details of these tests are provided elsewhere6.Figure 16 shows the experimental interaction diagramin comparison with those of RC columns of the samesize but with reinforcement ratios between 1 and 6%. Itis concluded that the hybrid column has an enhanced

behavior and is comparable to a 5% reinforced section.However, its flexural capacity is more comparable tothat of a 1% reinforced section.

Conclusions

A new hybrid column is proposed in the form ofconcrete-filled FRP tubes. For the section to work as acomposite unit, there is a need for mechanical bondbetween concrete and the tube. This can be achievedby providing shear connector ribs on the interior sur-face of the tubes. Design and manufacture of a specialmandrel and fabrication of the ribbed tubes were pre-sented. A total of 16 short beams and five long beam-columns made of concrete-filled FRP tubes were tested.The following conclusions were made:

1. Although shear demands for specimens of Series 1and 2 were quite different, it was clear that pres-ence of shear ribs could arrest any slippage orseparation and allow full utilization of the tube as acontinuous array of shear stirrups. Moreover,transfer of load between the tube and the concretecore was very effective. This was clear from theuniform distribution and propagation of cracks inthe tubes prior to failure.

2. Failure of concrete-filled tubes was not sudden norbrittle, as considerable deflections were observed.


Figure 14 Moment]curvature for specimens of Series 2

Figure 15 Test setup schematic for hybrid columns under axial]flexural loads

Figure 16 Interaction diagrams of hybrid column vs. RC columns

Ž .Ductility or pseudo-ductility was clear for bothshort and long beams under axial load, pure flex-ure, or a combination of the two.

3. The concept of hybrid columns with no internalreinforcement is feasible, as the experimental in-teraction diagram of tested columns correspondedto a reinforced concrete column of the same cross-section with more than 5% steel reinforcement.

Acknowledgements

This study was conducted with funding from theFlorida and the US Departments of Transportationunder Contract No. B-9895. The authors are grateful toMr. David Parks of the Marine Muffler Corp. forgenerously providing test materials and allowing theuse of their filament winding machine for fabrication ofthe specimens. The opinions and findings expressedhere, however, are those of the authors alone and notnecessarily the views of sponsoring agencies.


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13 Chajes, M. J., Januszka, T. F., Mertz, D. R., Thomson, T. A. andFinch, W. W., Shear strengthening of reinforced concrete beamsusing externally applied composite fabrics. ACI Structural Jour-nal, 1995, 92, 295]303

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design, manufacture and testing of a new hybrid column

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