Construction and Building Materials 52 (2014) 209–220

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Construction and Building Materials

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Low-grade glued laminated timber beams reinforced using improvedarrangements of bonded-in GFRP rods

0950-0618/$ - see front matter � 2013 Elsevier Ltd. All rights reserved.http://dx.doi.org/10.1016/j.conbuildmat.2013.11.044

⇑ Corresponding author. Tel.: +64 9 923 8536.E-mail address: G.Raftery@auckland.ac.nz (G.M. Raftery).

Gary M. Raftery a,⇑, Conor Whelan b

a Department of Civil and Environmental Engineering, Faculty of Engineering, The University of Auckland, New Zealandb Civil Engineering, College of Engineering and Informatics, National University of Ireland, Galway, Ireland

h i g h l i g h t s

� Low-grade glulam is reinforced using bonded-in GFRP rods.� The geometrical arrangement of the routed out grooves is of high importance.� The use of a number of smaller diameter rods to improve the bond surface area is of no advantage.� Enhancements for stiffness and ultimate moment capacity are reported on.� Improved utilisation of the compressive behaviour of the timber is obtained as demonstrated by strain profiles.

a r t i c l e i n f o

Article history:Received 29 July 2013Received in revised form 5 November 2013Accepted 12 November 2013Available online 7 December 2013

Keywords:Low-grade timberFibre reinforced polymersRod reinforcementGeometrical arrangementsBending test

a b s t r a c t

This paper describes an experimental test programme which examines the strengthening in flexure oflow-grade glued laminated timber beams (glulam) using bonded-in glass fibre reinforced polymer (GFRP)rods. Reinforcement configurations and arrangements previously not studied in the literature arereported on. The performance of unreinforced glulam beams, strengthened glulam beams using varyinggroove arrangements and different GFRP sized rods with single and double reinforcement was examined.The test results indicate that the geometrical arrangement of the routed out grooves is important andenhancements in the mechanical performance of the reinforced beams can be improved by reducingthe effects of stress concentrations. The use of a number of smaller diameter rods per groove to increasethe bond surface area proved to be of no advantage when reinforcing the beams. The unreinforcedlow-grade glulam exhibited linear elastic behaviour with brittle tensile failures in comparison to the non-linear behaviour of the beams strengthened on their tensile face. Using a reasonable reinforcement per-centage of 1.4% in the tension zone with circular routed grooves, mean stiffness enhancements of 11.2%and 13.9% for the global and local measurements were achieved respectively and a mean improvement inthe ultimate moment capacity of 68% was achieved in comparison to the unreinforced glulam beams.

� 2013 Elsevier Ltd. All rights reserved.

1. Introduction

In recent times, increasing focus is being placed on the develop-ment and promotion of sustainable materials in the constructionindustry [1]. Considerable emphasis is especially being placed onthe development of new engineered wood products. Timberpossesses many other advantages as a structural material in thatit is a natural renewable resource with a secure supply, recyclable,cost competitive, strong and is aesthetically pleasing. A class ofattractive materials that can offer many benefits to timber as areinforcement for both strengthening and rehabilitationapplications is that of fibre reinforced polymers (FRPs). Thesematerials are characterised by an excellent strength to weight ratio

when compared to other construction materials. The use of FRPmaterials for the strengthening of structural members in civilinfrastructure is effective and is reported on in the literature[2,3]. Good durability and fatigue characteristics are also portrayedby these materials in comparison to conventional constructionmaterials [4]. As a result maintenance costs are reduced with theiruse. To date, limited studies have been carried out on the reinforce-ment of laminated low-grade structural timber using internal rein-forcement. FRP materials are more suitable for the reinforcementof timber in comparison to steel because of the susceptibility ofsteel to corrosion when in contact with the moisture content ofthe timber. Fibre reinforced polymer materials can be easily pul-truded in various profiles. In contrast to the use of plate profiles,bars or rods adhesively bonded into routed out grooves in thestructural element preserve the architectural attractive nature ofwood as the reinforcement material is concealed. The glulam

210 G.M. Raftery, C. Whelan / Construction and Building Materials 52 (2014) 209–220

manufacturing process can also easily and effectively incorporatethe placement of the reinforcement and because the reinforcementis internally positioned, no additional depth is added to the beam.Mechanical protection is provided to the reinforcement by the sur-rounding wood material. The likelihood of premature delaminationis significantly reduced when the reinforcement is placed inter-nally as there is a much greater bond area. Glass fibre reinforcedplastics (GFRP) which incorporate E-glass are associated with lowcost and good mechanical properties and seem to be the mostappropriate fibre type for reinforcement of timber elements.

1.1. Adhesively bonded-in rod reinforcement in structural timberelements

The jointing between the dissimilar materials must be consid-ered when investigating the development of a reinforced gluedlaminated timber composite beam. As the timber is a natural mate-rial, the temperature and moisture content of the surroundingenvironment can change it physically, and this can put increasedstress on the connection between it and a dissimilar materialwhich may not swell and shrink as much. Adhesive bonding is aneffective technique to ensure that stresses between the reinforce-ment and the timber are transferred uniformly [5,6]. Stress concen-trations that are associated with mechanical fastening proceduresare eliminated [7]. Several experimental investigations have beenundertaken to examine the effect of modifying various parametersin relation to the quality of the bond formed in bonded-in connec-tions. Technical guidance is available in the literature on the use ofepoxy adhesives in the repair of structural glued laminated timber[8]. The use of epoxies alone is not recommended for the repair andstrengthening of wood stressed in tension parallel to grain. Carefulsurface preparation and thick adhesive bond lines to dissipate thestresses were key findings which were recommended from anexperimental test programme [9]. The pull-out behaviour of steelrods from timber was investigated and the strength was seen toimprove linearly with increasing bond area [10]. Epoxy adhesives,which possess gap filling properties, were reported to form a bondof high quality in comparison to other adhesive types. In anotherstudy it was found that both the strength and in particular, the fail-ure modes of the various specimen combinations were influencedby the moisture content of the timber at the time of bonding [11].Bainbridge et al. [12] reported that different adhesive types exhib-ited different failure modes in a programme of pullout tests. Thepullout capacity was seen to increase as embedment length andgroove size increased [13]. Contact cleaning methods to removeresidues were more effective than pressurized-air and better clean-ing techniques were found to generally increase pullout capacity.While it was reported that resin injection does enhance the struc-tural performance of a timber connection, widespread use was un-likely given the difficulties in fabrication [14]. It was determinedfrom a programme of pull-out tests that the bond capacity ismainly decided by the bond and/or the wood strength [15]. Fur-thermore, it was reported that savings in adhesive consumptioncould be achieved with the use of circular grooves for the proposedconnection but had little influence on the strength. In a studywhich examined surface preparations, it was reported that non-sanded bars were reported to be associated with slightly higherbond strengths than sanded bars and the bond strength is higherfor rods perpendicular to the grain than parallel to the grain [16].Experimental testing has shown that increasing adhesive bondlines thicknesses from 2 mm to 4 mm can achieve similar or higherstatic strengths but the fatigue strength of the joint is decreased[17].The use of creosote to treat the timber was shown to adverselyaffect the bond strength at the GFRP-timber interface [18]. Theshear strength of the samples without creosote treatment was28% higher than the treated samples. Raftery et al. [19,20] con-

cluded from extensive experimental test programmes that not onlydid the quality of the bond depend on the adhesive under exami-nation but it also depended on the FRP type that was being bonded.The use of certain conventional wood laminating adhesives andthin epoxy bond lines with a bond line thickness of 0.5 mm, wereshown to form strong durable bonds for specific combinations ofFRP plate material. A jointing system which used glued in steelrods between two timber beams was experimentally tested and anumerical model was developed which showed good agreementwith the experimental results [21]. In a programme of pull-outtesting tests, it was seen that a rougher external surface of therod provided a better bond performance [22]. Pull-out testing ofbasalt fibre reinforced polymer rods loaded perpendicular to thegrain was undertaken where it was found that the pull-out capac-ity increased almost linearly with the bonded length up to a max-imum value [23].

1.2. Strengthening of structural timber elements using FRP rod and barreinforcement

A comprehensive review of early efforts involving the reinforce-ment of timber elements can also be found in the literature [24].Experimental work undertaken using plate reinforcement and theuse of such reinforcement for the upgrade of low-grade laminatedfast-grown spruce is discussed in the literature [25]. The develop-ment of a nonlinear numerical model to predict the mechanicalbehaviour of reinforced low-grade laminated spruce is also de-tailed in a previous study [26]. Limited work is available in the lit-erature on the use of internal reinforcement for reinforcing low-grade fast grown wood. Gentille et al. [27] showed that the GFRPbars enhanced the bending strength of creosote treated Douglasfir beams, overcame the effects of local defects and changed thefailure mode from brittle tension to compression failure. Svecovaand Eden [28] reported that GFRP bars bonded into timber beamsboth as flexural and shear reinforcement succeeded in increasingthe ultimate strength and also succeeded in reducing the variabil-ity of the members. A composite system to join glulam beams to-gether using CFRP rods was investigated as an alternative to steelplates and bolts and compared well mechanically to monolithicallytested beams [29]. Johnsson et al. investigated the use of bonded inbars for the reinforcement of glued laminated timber elements andproposed an analytical model was proposed to predict the anchor-age length [30]. Schober and Rautenstrauch examined the use ofembedded reinforcement for solid timber and concluded that thewood, CFRP and bonding agent had a considerable influence onthe overall strength of the specimen [31]. Ribeiro et al. reportedthat the position of the reinforcement in the cross section shouldbe optimised. Placement of the reinforcement at the neutral axiswas not recommended [32]. Jankowski et al. tested hundred yearold beams and determined the strain distributions using electricalresistance strain gauges along with the photo-elastic coating tech-nique [33]. It was concluded that the effectiveness of the reinforce-ment significantly depended on the quality of the wood-FRP bondand that the reinforcement was useful for restoration purposes.The use of steel inserts for reinforcement in historic buildingshas achieved good results [34]. Gentry examined the use of epoxybonded-in FRP pins to increase the shear capacity of glulam. Thestrategy showed potential economic and structural advantagesespecially when the FRP was coupled with low-grade, low valuefinger jointed timber [35]. Nowak et al. examined the use of inter-nal CFRP reinforcement to reinforce pine-wood beams with defectsand achieved noteworthy enhancements in load-carrying capacity[36]. An economic analysis which examined the use of FRP rein-forcement for glued laminated timber manufactured using woodsourced in the north east of the United States demonstrated thatproduction cost advantages were limited to deeper beams which

G.M. Raftery, C. Whelan / Construction and Building Materials 52 (2014) 209–220 211

required a high bending strength capacity [37]. Martin et al.anticipated that commercial success of reinforced wood waspossible because of the significant enhancements that could beachieved with the use of less desired, more easily available woodmaterial [38].

1.3. Objectives of the present study

The objective of this research is to examine the strengthening oflow-grade glued laminated timber using bonded-in pultrudedglass fibre reinforced polymer rods. The experimental test pro-gramme involved the fabrication and testing in flexure of unrein-forced and strengthened glued laminated timber beams whichwere manufactured from low-grade spruce. The effects of contrast-ing reinforcement and groove arrangements which have not previ-ously been examined in the literature are studied. The mechanicalperformance of the reinforced beams is compared with that of theunreinforced glulam beams with regard to the load–deflectionbehaviour, failure mode, enhancements in both stiffness and ulti-mate moment capacity as well as the changes in strain profiledistribution.

2. Materials

2.1. Timber

The timber species studied in this experimental test programme was Irishgrown Sitka spruce. This species is very fast grown. The timber used in the studywas sourced from the same sawmill so that variability resulting from contrastingenvironmental growth conditions was reduced. The timber was C16 grade in accor-dance with EN 338 [39], plain sawn cut and was kiln dried in the sawmill to approx-imately 18% moisture content. The boards were 4200 mm in length with a nominalsection size of 96 mm � 44 mm. The boards were stored for three months in a con-ditioned environment of 65 ± 5% relative humidity and 20 ± 2 �C temperature oncedelivered to the laboratory. A mean equilibrium moisture content of 12.4% withstandard deviation of 0.7% was obtained after the conditioning period and a meanboard density of 385 kg/m3 with a standard deviation of 39 kg/m3 was recorded.

2.2. FRP

The GFRP rods used as the reinforcement were Aslan 100. This FRP material iscorrosive resistant, non-conductive and weighs approximately one fourth that ofsteel reinforcing rods. This FRP material is manufactured from continuously drawnE-glass roving saturated with vinyl ester resin. The surface of the rods are purposelydeformed and are characterised by a helical wrapped nature which possesses a sandcoating in order to enhance bonding when using epoxy adhesives. The mechanicalproperties of the FRP rods in comparison to Irish grown Sitka spruce are shown inTable 1.

2.3. Adhesive

The adhesives that were selected for bonding the timber laminations and forbonding the rods into the beams were examined in previous research programmes[19,20,44]. From the results obtained during this testing, a phenol resorcinol form-aldehyde (PRF) was selected for the bonding of the wood-wood laminations. PRFwas also more easily sourced in Ireland in comparison to melamine urea formalde-hyde. Also from the results, the epoxy adhesive, Sikadur 31, a well-recognised civilengineering two-component thixotropic adhesive, showed favourable performancefor both ambient and accelerated ageing conditions between FRP and wood.

Table 1Mechanical properties of FRP and timber.

Material Ultimate tensile strength(N/mm2)

Modulus of elasticity in tension(N/mm2)

GFRP rods 620a 45,000a

47,000–52,000b

Irish grown Sitkaspruce

23.7c 8111c

a GFRP rod material properties as reported by manufacturer [40].b GFRP rod material properties as tested [41].c In-grade testing with mean moisture content of 12% and mean density of

403 kg/m3 [42,43].

3. Experimental testing

3.1. Test programme

The test programme involved the fabrication and testing ofunreinforced low-grade glulam beams and reinforced low-gradeglulam beams. Stiffness testing and testing to failure wasundertaken for each beam. The beam configurations are shownin Fig. 1 and the test programme is shown in Table 2. Alldimensions shown in the figures are in millimetres. Five repeti-tions were tested for each configuration in order to ensure confi-dence in the results. The unreinforced beams were tested inorder to act as a control for the reinforced beams. Three reinforce-ment strategies with tension only reinforcement were studiedwhile the fourth reinforcement strategy included both tensionand compression reinforcement. The reinforcement percentagesstudied were as follows: 1.05% reinforcement (Phase B); 1.4%reinforcement (Phase C); 1.4% reinforcement (Phase D) and 2.8%reinforcement (Phase E).

3.2. Beam fabrication

The initial ranking of the lamination stock was undertaken basedon the results from a three point bending mechanical stress grader[45]. Visual grading with reference to IS 127 was also undertakenfor each of the boards [46]. Boards that were associated with bow,spring or cupping that was considered to be significant, were re-moved from the experimental test stock. Boards in which the knotarea ratios, whether margin or total, or other strength reducing de-fects such as fissures were considered significant, were used only inthe internal compression laminations, internal tension laminationsor beam core so that premature failure of the beams would not occurif such boards were used in the most extremely stressed tension andcompression laminations. Each beam comprised five laminations of38 mm thickness after knife-planing, giving an overall beam depthof 190 mm. The testing of the beam was to be undertaken using afour point bending test arrangement with a span-depth ratio of18:1 [47]. This meant a span of 3420 mm was used. Each boardwas cut from 4200 mm to the required lamination length fromeither or both sides. This procedure allowed for the quality of theboards to be maximised in each lamination and that defects weremoved away from the zone of the maximum bending moment areaand hence premature failure was avoided. The manufacture of thebeams involved optimising the performance of the beams such thatthe best quality laminations were used as the most extremelystressed tension laminations at the bottom of each beam. The rank-ing procedure involved placing the best quality lamination deter-mined from the grading procedure as the most highly stressedtension lamination at the soffit of Beam 1 with the next best qualitylamination being placed at the bottom of Beam 2. This methodicalapproach was used as the beam number ascended. The arrangementof laminations in the critical compressive zone at the top of thebeams followed a similar procedure. The secondary tension, second-ary compressive and core laminations were all allocated using theranking procedure detailed above. As a result of this approach lowerbeam numbers would be theoretically stronger than higher beamnumbers. Bonding of the laminations was undertaken using a PRFadhesive with no more than two hours being permitted to elapse be-tween the planing process and application of the adhesive. Theadhesive was spread rate at a rate of 400 g/m2. The laminations wereoriented such that the pith was towards the top side of the board asspecified in EN 386 [48]. Pressure was applied to the closed assemblyat a rate of 0.7 N/mm2 for 24 hours in an environment of 65 ± 5% rel-ative humidity and 20 ± 2 �C temperature. Placing of the GFRP rein-forcement involved routing of the square or circular shaped grooves

Fig. 1. Beam configurations: (a) Phase A. (b) Phase B. (c) Phase C. (d) Phase D. (e) Phase E.

Table 2Experimental test programme.

Phase Description Repetitions Beam numbers

A Unreinforced 5 1, 8, 15, 22, 29B Single reinforced with 6 mm rods in square grooves 5 2, 9, 16, 23, 30C Single reinforced with 12 mm rods in square grooves 5 3, 10, 17, 24, 31D Single reinforced with 12 mm rods in circular grooves 5 4, 11, 18, 25, 32E Double reinforced with 12 mm rods in circular grooves 5 5, 12, 19, 26, 33

212 G.M. Raftery, C. Whelan / Construction and Building Materials 52 (2014) 209–220

using a computer numerically controlled machine. All the groovesfacilitated a 2 mm bondline around the reinforcement. Because ofthe inherently coarse surface of the GFRP material, the materialwas simply wiped clean with methylated spirits before the epoxyadhesive was applied. The epoxy adhesive was applied within twohours of placing the reinforcement. A 2 mm thick wire was wrappedaround the reinforcement rod at 400 mm intervals which allowed auniform adhesive thickness to surround the reinforcement. Thebeams were subsequently placed in an environment of 65 ± 5%relative humidity and 20 ± 2 �C temperature for a period of at least30 days to ensure the adhesive was fully cured.

3.3. Beam testing

Testing of the beams was carried out using the four point bend-ing experimental arrangement as stated in EN 408 [47]. The testsetup is shown in Fig. 2. Testing was initially carried out on allthe beams in their unreinforced state for both global and localstiffness using a 500 kN hydraulic actuator. The global stiffnessmeasurement is recorded by means of an inverted linear variabledifferential transformer (LVDT) which is located at midspan. Thereading of deflection is taken relative to the supports andhence is over the entire span of the beam. The local stiffness

G.M. Raftery, C. Whelan / Construction and Building Materials 52 (2014) 209–220 213

measurement is restricted to the zone of maximum bending mo-ment where no shear effects are present. A second LVDT is posi-tioned in a hanger which is suspended from the neutral axis overa distance of five times the depth of the beam (Fig. 2). The localstiffness measurement also excludes the effects of indentationsin the timber at the supports. A constant crosshead displacementrate of 0.57 mm/s was used to load the beams after a preload of300 N was applied before the test commenced. Local indentationsfrom the loading heads and supports were minimised by placingsteel plates of dimensions 95 mm � 90 mm � 10 mm at these loca-tions. At no stage during the test, did the load exceed the elasticlimit or 40% of the ultimate load. The beams were laterally re-strained at positions approximately 300 mm outside of the loadingheads and frictional effects were reduced to a minimum by the useof polytetrafluoroethylene (PTFE) strips sliding over each other.The experimental arrangement, shown in Fig. 2, was also employedfor testing of the beams when the reinforcement was added. How-ever, for the beams which included tension rod reinforcement only,the hanger to accommodate the LVDT for the measurement of thelocal stiffness in these beams was at a calculated lower depth.

Before loading to failure for any of the beams, the local stiffnessLVDT and hanger were removed. The time to failure during thetesting was 300 ± 120 s [37]. The strain distribution behaviourwas recorded on Beam 16 (Phase B), Beam 10 (Phase C), Beam 24(Phase C) and Beam 19 (Phase E). Beam 16 and Beam 19 were se-lected as they were in the midrange of the beams manufacturedand were most representative of the beams in the test programme.Biased results may have been achieved if the weakest or strongestbeams were instrumented. Two beams for Phase C were examinedto see if any major deviations existed between beams in the samephase. Theoretically, Beam 10 should be stiffer and stronger thanBeam 24 as a result of the selection procedure for the laminatesduring manufacturing. No unreinforced beam was instrumentedas the strain behaviour of an unreinforced low-grade glued lami-nated timber beam was recorded in a previous research pro-gramme [25]. No strain gauges were placed on any Phase Dbeam because of cost and time constraints. The instrumentationinvolved positioning strain gauges of 60 mm gauge length (PL-60-11) at midspan throughout the depth of each beam, on bothfaces of the timber laminations. Strain gauges were applied on bothsides of the beam in order to account for twist that might occurduring flexure. The electrical resistance of the gauges was 120 X.The gauges were placed at mid-depth in the lamination and a long-er gauge of 60 mm in length was selected as this allowed a meanestimation of the stresses to be obtained rather than a spot value.The arrangement used for the strain gauges as shown for a Phase Cbeam is shown in Fig. 3.

4. Experimental results and discussion

4.1. Load–deflection behaviour

4.1.1. Load–deflection behaviour, Phase A – Unreinforced beamsThe load–deflection behaviour to failure of the unreinforced

control beams are shown in Fig. 4. All the beam exhibited linear

Fig. 2. Stiffness tes

elastic behaviour until their point of failure. The beams failedthrough tensile failure of the bottom lamination with the failurecracks emanating from defects or irregularities in the wood. Nocompression wrinkling was noted in the compression zone ofany of the beams. This was despite the fabrication procedure whichwas followed during manufacturing of the glulam beams wherebythe best quality laminations were placed in the tension zone andthe best effort was made to move defects away from the zone ofthe maximum bending moment during the four point bending test.However, because of the low-grade of the timber that was beingexamined in this study, significant defects were still present inthe bottom tension lamination. For timber with defects, the aver-age tensile strength is lower than the average compressivestrength because of the much greater influence of the presenceof defects on the tensile strength [38]. Therefore, the yield stressin tension is reached in these beams before the yield stress in com-pression is reached and hence no nonlinear behaviour is experi-enced. A typical beam failure is shown in Fig. 5.

4.1.2. Load–deflection behaviour, Phase B – Tension reinforcementusing multiple 6 mm diameter rods

As can be seen from the load-deflection curves shown in Fig. 6,non-linear behaviour can be introduced into the beams when theGFRP reinforcement is strategically located at the bottom of thebeams. The amount of ductility depends on the quality of the tim-ber that is present in the bottom lamination after the reinforce-ment is bonded into the routed out grooves as some defectswere removed during this procedure and replaced with the epoxybonded in GFRP rods. Failure in the beams resulted from irregular-ities such as defects or slope of grain around knots as can be seen inFig. 7. Cracking of the adhesive only occurred after the significantcracks in the timber had initiated. Compression buckling in thetop lamination which contributed to the nonlinear behaviour wasparticularly visible in Beams 2 and 16 of the reinforced beams ascan be seen in Fig. 8. Beam 2 achieved a noticeably higher ultimateload carrying capacity and this is as a result of a higher quality lam-ination being used in the manufacture of this beam as discussedpreviously. No premature debonding of the reinforcement was re-corded for any of the beams during testing.

4.1.3. Load–deflection behaviour, Phase C – Tension reinforcementusing 12 mm diameter rods in square grooves

The load–deflection behaviour from the Phase C beams isshown in Fig. 9. An increase in load carrying capacity is noted aswell as increased consistency in comparison to the Phase B beams.Depending on the bottom timber lamination nonlinear behaviouris also associated with these beams prior to failure. Fig. 10 showsa typical compression wrinkle that was experienced in thesebeams. Again, all beams failed as a result of excessive tensile stres-ses in the bottom laminate with failure most commonly initiatingat knots as shown in Fig. 11. It should be also noted that the PhaseC beams are not only slightly stronger but the reinforcement ismuch easier to install. Firstly, only two grooves were routed intothe tension face of the beam and also only two rods need to bebonded in, whereas for the Phase B specimens three grooves are

t arrangement.

Fig. 3. Strain gauges arrangement for Phase C beam.

Fig. 4. Load vs. deflection behaviour for unreinforced beams (Phase A).

Fig. 5. Tension failure at defect in Beam 22 (Phase A).

Fig. 6. Load vs. deflection behaviour for reinforced beams using 6 mm diameterrods in multiple grooves on the tension face (Phase B).

Fig. 7. Tension failure near irregularity in Beam 2 (Phase B).

Fig. 8. Compression buckling in Beam 2 (Phase B).

214 G.M. Raftery, C. Whelan / Construction and Building Materials 52 (2014) 209–220

routed out and six rods are bonded in. The epoxy adhesive showedno signs of premature fracture and only cracked after the timber inthe beams was severely ruptured.

4.1.4. Load–deflection behaviour, Phase D – Tension reinforcementusing 12 mm diameter rods in circular grooves

In comparison with the behaviour load deflection behaviourdemonstrated by the Phase C beams, there is a significant improve-ment in the load carrying capacities of the Phase D beams, as seenin Fig. 12. The only difference between these phases, apart from thevariability in the timber, is the shape of the routed groove, which issquare in Phase C and circular in Phase D. The approach to limit the

effects of the variability in the timber was already explained. It isalso important to note that the Phase C beams were manufacturedwith laminations of marginally higher quality than the Phase Dbeams and yet the Phase D beams, using the circular routedgrooves appear to outperform the Phase C beams which includedthe reinforcement placed in the square routed grooves. Therefore,it is adjudged that by using the circular routing head, stress con-centrations in the groove are significantly reduced. These resultssuggest that undertaking such a procedure can further increasethe load carrying capacity when using bonded-in reinforcementin low-grade glued laminated beams. All of the beams failed by

Fig. 9. Load vs. deflection behaviour for reinforced beams using 12 mm diameterrods in rectangular routed out grooves on the tension face (Phase C).

Fig. 10. Significant compression buckling in Beam 24 (Phase C).

Fig. 12. Load vs. deflection behaviour for reinforced beams using 12 mm diameterrods in circular routed out grooves on the tension face (Phase D).

Fig. 13. Load vs. deflection behaviour for reinforced beams using double reinforce-ment (Phase E).

G.M. Raftery, C. Whelan / Construction and Building Materials 52 (2014) 209–220 215

means of a tensile failure in the bottom lamination and all showedsigns of compression wrinkles in the top lamination. Reserve loadcarrying capacity was demonstrated by Beams 25 and 32 after theinitial fracture occurred in the bottom tension lamination.

4.1.5. Load–deflection behaviour, Phase E – Tension and compressionreinforcement using 12 mm diameter rods in circular grooves

The load deflection behaviour from the Phase E beams whichincluded compression reinforcement as well as tension reinforce-ment can be seen in Fig. 13. The behaviour that was predominatelyexperienced by these beams is linear elastic to failure. In these

Fig. 11. Tensile fracture by defect in bottom lamination (Phase C).

beams, the neutral axis was approximately positioned at mid-depth because of the balanced reinforcement arrangement thatwas adopted as well as the methodology that was used to con-struct the beams. As well as acting as a reinforcement, the GFRPmaterial that is placed in the compression zone of the beam assistsin reducing the knot area ratio in the top lamination, and hence re-duces the likelihood of plastic yielding occurring in the timber.Therefore, the ultimate tensile strength of the timber is generallyreached in the bottom lamination prior to the ultimate compres-sive strength being reached in the top lamination. No visible signsof compression buckling were evident in any of the beamsalthough limited ductility is present in Beam 5. This was becausethe bottom lamination used in this beam was of a higher qualitythan the bottom lamination used in the other beams in the beamphase. No buckling of the compression GFRP reinforcement wasnoted in any of the beams and there was no problems with thebond of the compression reinforcement. All the beams failed atknots in the tension zone which were located towards one of theedge areas. With the exception of Beam 26, all the beams failureswere catastrophic in nature after initial audible indications ofcracking as shown in Fig. 14. Beam 26 failed at a lower load thanall the other beams as a result of a significant knot in the margintensile zone but yet maintained some reserve capacity as the ten-sion reinforcement remained in place.

4.2. Stiffness testing

4.2.1. Stiffness testing, Phase A – Unreinforced beamsThe unreinforced Phase A beams in this test programme act as

control reference beams and were only tested in their unreinforced

Fig. 14. Typical catastrophic failure of Phase E beams (Beam 12).

Fig. 15. Stiffness testing: Phase A. Key: A = Global stiffness unreinforced beams;B = Local stiffness for unreinforced beams.

Fig. 16. Stiffness testing: Phase B. Key: A = Global stiffness for beams whenunreinforced; B = Local stiffness for beams when unreinforced; C = Global stiffnessfor 6 mm rods as tension reinforcement in multiple grooves; D = Local stiffness for6 mm rods as tension reinforcement in multiple grooves.

Fig. 17. Stiffness testing: Phase C. Key: A = Global stiffness for beams whenunreinforced; B = Local stiffness for beams when unreinforced; E = Global stiffnessfor 12 mm rods as tension reinforcement in square grooves; F = Local stiffness for12 mm rods as tension reinforcement in square grooves.

216 G.M. Raftery, C. Whelan / Construction and Building Materials 52 (2014) 209–220

state prior to testing to failure. A mean global stiffness of 4.85� 1011 N mm2 with standard deviation of 0.1 � 1011 N mm2 anda mean local stiffness of 5.16 � 1011 N mm2 with standard devia-tion of 0.19 � 1011 N mm2 was achieved as shown in Fig. 15. Thereare a number of reasons for the deviation in magnitude betweenthe global and local stiffness readings. Shear deformation is presentin the global stiffness test while no shear deformation is present inthe local stiffness test. There is a possibility of indentation at thesupports as a result of the low density of the wood that is usedin the manufacturing of the beams. This can consequently lowerthe magnitude of the global stiffness test results. The mean globalstiffness for the Phase A beams is within 3.5% of the mean globalstiffness value of 5.01 � 1011 N mm2 which was determined forall the beams in the test programme when tested in their unrein-forced state. A standard deviation of 0.33 � 1011 N mm2 was asso-ciated with this mean value.

4.2.2. Stiffness testing, Phase B – Tension reinforcement using multiple6 mm diameter rods

The stiffness of the Phase B beams was measured twice. They weretested, first of all, after the curing of the glulam beams and secondly,after the addition of the GFRP rods. The results of the tests are shownin Fig. 16. In their unreinforced state, the Phase B beams had a meanglobal stiffness of 5.03� 1011 N mm2 with a standard deviation of0.42� 1011 N mm2 and a mean local stiffness of 5.91 � 1011 N mm2

with a standard deviation of 0.66 � 1011 N mm2. In their reinforcedstate, the Phase B beams had a mean global stiffness of5.36 � 1011 N mm2 with a standard deviation of 0.31 � 1011 N mm2

and a mean local stiffness of 6.41� 1011 N mm2 with a standarddeviation of 0.78 � 1011 N mm2. This represents a mean globalstiffness increase of 6.7% and a mean local stiffness increase of 8.2%.

The local stiffness results are of a higher magnitude than the globalstiffness results as explained for the Phase A beams. However, theindentation at the supports would be reduced because of the presenceof the GFRP rods epoxy bonded into the soffit.

4.2.3. Stiffness testing, Phase C – Tension reinforcement using 12 mmdiameter rods in square grooves

The Phase C beams were tested both in their unreinforced andreinforced states. The results are shown in Fig. 17. In their unrein-forced state, the Phase C beams achieved a mean global stiffness of4.79 � 1011 N mm2 with a standard deviation of 0.4 � 1011 N mm2

and a mean local stiffness of 5.32 � 1011 N mm2 with a standarddeviation of 0.44 � 1011 N mm2. In their reinforced state, the PhaseC beams had a mean global stiffness of 5.29 � 1011 N mm2 with astandard deviation of 0.39 � 1011 N mm2 and a mean local stiffnessof 5.85 � 1011 N mm2 with a standard deviation of 0.42 �1011 N mm2. This corresponds to a mean global stiffness increaseof 9.5% and a mean local stiffness increase of 10.8%. These val-ues represent a larger increase in stiffness in comparison to thePhase B orientation, due to the increased percentage ofreinforcement.

4.2.4. Stiffness testing, Phase D – Tension reinforcement using 12 mmdiameter rods in circular grooves

The stiffness results of the Phase D beams are shown in Fig. 18.The beams were tested in their unreinforced and reinforced statefor both local and global stiffness. In their unreinforced state, thePhase D beams had a mean global stiffness of 5.22 � 1011 N mm2

with a standard deviation of 0.32 � 1011 N mm2 and a mean local

Fig. 18. Stiffness testing: Phase D. Key: A = Global stiffness for beams whenunreinforced; B = Local stiffness for beams when unreinforced; G = Global stiffnessfor 12 mm rods as tension reinforcement in circular grooves; H = Local stiffness for12 mm rods as tension reinforcement in circular grooves.

G.M. Raftery, C. Whelan / Construction and Building Materials 52 (2014) 209–220 217

stiffness of 6.05 � 1011 N mm2 with a standard deviation of0.25 � 1011 N mm2. In their reinforced state, the Phase D beamshad a mean global stiffness of 5.8 � 1011 N mm2 with a standarddeviation of 0.38 � 1011 N mm2 and a mean local stiffness of6.85 � 1011 N mm2 with a standard deviation of 0.35 � 1011

N mm2. This represented a mean increase of 11.2% and 13.9% inthe global stiffness and local stiffness of the reinforced beam,respectively and exceeds the performance increase demonstratedby the Phase C beams. The reinforcement percentage used inthe Phase D beams was identical to the percentage used in thePhase C beams. The only difference in the manufacturing proce-dure was the shape of the groove in which the reinforcementwas placed.

4.2.5. Stiffness testing, Phase E – Tension and compressionreinforcement using 12 mm diameter rods in circular grooves

The Phase E beams were stiffness tested after fabrication and againafter reinforcement had been inserted on the tension and compres-sion faces of the beams. The results are presented in Fig. 19. In theirunreinforced state, the Phase E beams had a mean global stiffness of5.15 � 1011 N mm2 with a standard deviation of 0.23 � 1011 N mm2

and a mean local stiffness of 5.68 � 1011 N mm2 with a standard devi-ation of 0.502 � 1011 N mm2. In their reinforced state, the Phase Ebeams had a mean global stiffness of 6.28� 1011 N mm2 with a stan-dard deviation of 0.15� 1011 N mm2 and a mean local stiffness of7.33 � 1011 N mm2 with a standard deviation of 0.3 � 1011 N mm2.This corresponds to an increase in the mean global stiffness of 22%and an increase in the mean local stiffness of 29.6%. The additionalreinforcement added in the compression zone therefore contributes

Fig. 19. Load-deflection behaviour for unreinforced beams (Phase E). Key:A = Global stiffness for beams when unreinforced; B = Local stiffness for beamswhen unreinforced; J = Global stiffness for double reinforced beams; and K = Localstiffness for double reinforced beams.

to a significant improvement in the overall stiffness of the section incomparison to the other reinforcement configurations studied inthe test programme.

4.3. Ultimate moment capacity

The ultimate moment capacity Mult achieved by each of the dif-ferent orientations is shown in Table 3. The strengthening effectwhen incorporating the reinforcement is evident from the datapresented. Even with the least effective Phase B orientation that in-cluded the 6 mm diameter GFRP rods as tension reinforcement, anaverage increase of 39% with an average ultimate moment capacityof 26.62 kNm was achieved in comparison to 19.15 kNm for theunreinforced control specimens. The configuration employed inPhase C, where two 12 mm diameter rods were bonded intogrooves in the soffit of the beam achieved a greater enhancementin ultimate moment capacity in comparison to Phase B. Thesebeams achieved an average ultimate moment capacity of 27.35kNm which represented an increase of 43% in comparison to theunreinforced control beams. It should be noted that the reinforce-ment configuration used in the Phase C beams contributed to areinforcement percentage of 1.05% in comparison to 1.4% used inthe Phase B beams. However, a reduced bond surface area wasassociated with the Phase C reinforcement configuration but thishad no negative influence on the performance of the system. ThePhase D beams which included the 12 mm GFRP rods positionedin the circular routed grooves achieved an average ultimate mo-ment capacity of 32.18 kNm which represented an average in-crease of 68% in comparison to that achieved by the unreinforcedbeams. The primary difference between the Phases D and C orien-tations is the shape of the routing head which is deemed to be theinfluential factor in the increased performance exhibited by thePhase D beams. It is considered that stress concentrations are min-imised when using a circular shaped router head which can delaythe fracturing of the reinforced beam. Minor deviations betweenthe timber laminations of the beams would also be present. As ex-pected, Phase E which included both tension and compressionreinforcement and which had the greatest percentage reinforce-ment achieved the highest average ultimate moment capacity withan average value of 38.01 kNm which corresponds to an averageincrease of 98.5% over the control specimens. The deviation in glo-bal stiffness readings between the unreinforced Phase A beams andthe glulam beams for each of the beam phases when tested in theirunreinforced state were as follows: Phase B (3.7%); Phase C(�1.2%); Phase D (+7.5%); Phase E (+6.2%).

4.4. Strain profile distribution

The Moment vs. Strain behaviour for Beam 16 is shown inFig. 20. It is believed that Beam 16 was a good representation ofthe mechanical properties of the beams in Phase B as the lamina-tions used to fabricate this beam are close to average of themechanical properties of the laminations used to fabricate all ofthe beams which were reinforced with the smaller diameter rods.It can be seen that nonlinear strain behaviour is recorded prior tofailure. This is in contrast to linear strain behaviour to failure beingrecorded for instrumented unreinforced low-grade laminatedbeams in a previous research beam test programme [25]. The pres-ence of ductile compressive strain is considered favourable as it is awarning sign of the incipient failure of the structural member.However, the strain gauges positioned on the top face of the beam(SG1) and on the sides of the top lamination did not record themaximum strain (SG2, SG3) as the maximum compressive strainwas localised at a position of a knot offset from the midspan ofthe beam. Nonlinear behaviour was recorded on the gauges aslow as the centre lamination in the beam which shows the

Table 3Ultimate bending moment, Mult.

Phase A Mult (kNm) Phase B Mult (kNm) Phase C Mult (kNm) Phase D Mult (kNm) Phase E Mult (kNm)

Beam 1 22.94 Beam 2 30.32 Beam 3 26.59 Beam 4 34.03 Beam 5 44.37Beam 8 19.78 Beam 9 25.96 Beam 10 27.11 Beam 11 32.8 Beam 12 41.7Beam 15 17.45 Beam 16 26.01 Beam 17 29.5 Beam 18 34.09 Beam 19 35.88Beam 22 18.58 Beam 23 25.93 Beam 24 28.15 Beam 25 31.75 Beam 26 31.64Beam 29 17.01 Beam 30 24.88 Beam 31 25.42 Beam 32 28.22 Beam 33 36.48

Average (kNm) 19.15 – 26.62 – 27.35 – 32.18 – 38.01St dev (kNm) 2.38 – 2.12 – 1.55 – 2.42 – 5.04Median (kNm) 18.58 – 25.96 – 27.11 – 32.8 – 36.48COV 0.12 – 0.08 – 0.06 – 0.08 – 0.13Increase (%) – – 38.99 – 42.83 – 68 – 98.48

Fig. 20. Moment vs. Strain for Beam 16 (Phase B). Fig. 22. Moment vs. Strain for Beam 10 (Phase C).

218 G.M. Raftery, C. Whelan / Construction and Building Materials 52 (2014) 209–220

considerable effect of using the GFRP reinforcement with the low-grade timber. Strain gauge profiles from gauges on opposite sidesof each individual lamination should ideally overlap if the materialwas homogeneous and there was no twist in the beam. Segregationof readings at higher loads is believed to be due to the micro-crack-ing of the beam on the tension side during testing, and due to com-pression wrinkles developing in the compression zone of the beam.Strain behaviour in the tension zone was predominantly linearelastic to failure with the exception of some deviations as a resultof microcracking before failure.

The strain profile for various moments for Beam 21 is illustratedin Fig. 21. The reinforcement in the tension zone is seen to lowerthe location of the neutral axis when behaviour of the hybrid beamis in the elastic region (10 kNm, 20 kNm). As the applied momentincreases and ductility is introduced in the top laminations, the

Fig. 21. Strain profile for Beam 16 (Phase B).

neutral axis is seen to deepen in the section as can be seen fromthe strains recorded when a moment of approximately 26 kNm isapplied.

The Moment vs. Strain plots for the two instrumented Phase Cbeams, Beams 10 and 24, are shown in Figs. 22 and 23, respec-tively. No significant difference exists between the data recordedfor both beams. Similar to the instrumented Phase B beam, thestrain gauges on the top laminations do not record readings fromthe mostly highly strained localised zone as this occurred at adefect within the zone of maximum bending moment but awayfrom the midspan. In particular, the distortion of readings fromSG1 and SG2 for Beam 24 is due to the presence of a large compres-sion wrinkle away from the position of the gauges. Again, it is seenthat nonlinear behaviour is recorded by the gauges as low as thecentre lamination in the beam (SG6 and SG7). More significant

Fig. 23. Moment vs. Strain for Beam 24 (Phase C).

Fig. 24. Strain profile for Beam 10 (Phase C).

Fig. 25. Strain profile for Beam 24 (Phase C).

Fig. 27. Strain profile for Beam 19 (Phase E).

G.M. Raftery, C. Whelan / Construction and Building Materials 52 (2014) 209–220 219

nonlinear behaviour is recorded by these gauges on Beam 24 as thewood used in the manufacturing of this beam is of lower strengthand stiffness. It is also seen from the strain profiles in Figs. 24and 25 that the neutral axis is at a lower depth in the Phase Cbeams than the Phase B beam when behaviour is in the elasticregion. This is because of the higher percentage of reinforcementused in the Phase C beams in comparison to the Phase B beams.The position of the neutral axis appears to be at a margin-ally lower depth in Beam 24 in comparison to Beam 10 as thequality of the wood used to manufacture this beam is of a lowerstiffness. Strain behaviour to failure in the tension zone of bothbeams, with the exception of some micro-cracking, is largely seento be elastic.

Fig. 26. Moment vs. Strain for Beam 19 (Phase E).

The Moment vs. Strain plot for Beam 19 (Phase E), which in-cluded both compression and tension reinforcement is shown inFig. 26. The strain gauges on the top of the beam (SG1) recordshigher strains in comparison to the single reinforced beams. Thisis believed to be because by positioning reinforcement in the com-pression zone, defects such as knots are removed which lower thelikelihood of localised nonlinear behaviour occurring away fromthe strain gauge. The strain gauges, SG4 and SG5, show limitednonlinear behaviour in comparison to the single reinforced beamsas the neutral axis in the double reinforced section is closer to mid-depth of the beam as can be seen in the strain profile plot in Fig. 27.Furthermore, because of the overall stiffer section and the use of2.8% reinforcement in the double reinforced Phase E section incomparison to 1.05% (Phase B) and 1.4% (Phase C), the strains re-corded at the bottom of Beam 19 (Phase E) are considerably lowerwhen moments of 10 kNm and 20 kNm are applied in comparisonto strains recorded at the bottom of the single reinforced sectionfor the same applied moments. Hence, the double reinforced sec-tions demonstrate significantly greater load carrying capacities.Deviations in strain readings on gauges at the same depths oneither side of the beam are considered to be as a result of non-homogenous properties in the wood and the initiation of micro-cracks in the tension zone at higher moment levels.

5. Conclusions

The results of a test programme to reinforce low-grade gluedlaminated timber in flexure using bonded in GFRP rods were dis-cussed. Several important observations have been made as re-ported in the following:

� An enhanced performance in relation to improvements in stiff-ness and ultimate moment capacity was achieved with the useof circular routed out grooves in comparison to square groovesfor the reinforcement.� The use of a number of smaller diameter rods per groove to

increase the bond surface area proved to be of no advantagewhen reinforcing the low-grade glulam if the adhesive selectedforms a bond of adequate quality.� The unreinforced beams in the test programme failed in a linear

elastic brittle tensile manner in comparison to the ductilebehaviour of the reinforced beams which included tension rein-forcement only.� Strain profiles indicated that greater utilisation of the nonlinear

compressive characteristics of the timber are achieved whengreater percentages of reinforcement are utilised. Furthermore,increased nonlinear behaviour is associated with the reinforcedbeams when lower quality laminations are used.

220 G.M. Raftery, C. Whelan / Construction and Building Materials 52 (2014) 209–220

� The use of 1.4% reinforcement percentage strategically posi-tioned in circular shaped grooves in the tension zone resultedin mean stiffness enhancements of 11.2% and 13.9% for theglobal and local measurements and mean improvement in theultimate moment capacity of 68% in comparison to the unrein-forced glulam beams.� The use of 2.8% reinforcement, with 1.4% tension reinforcement

and 1.4% compression reinforcement, resulted in mean stiffnessenhancements of 22% and 29.6% for the global and local mea-surements and mean improvement in the ultimate momentcapacity of 98.5% in comparison to the unreinforced glulambeams.

Acknowledgments

The authors would like to thank Civil Engineering, National Uni-versity of Ireland, Galway for financially assisting the postgraduateresearch student, Conor Whelan. The research project was under-taken at National University of Ireland, Galway and the primaryauthor would like to sincerely thank the National University of Ire-land, Galway for financially supporting the research. The donationof the GFRP materials from Banagher Concrete is most appreciated.

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