Composites: Part B 42 (2011) Contents lists available at ScienceDirect. Composites: Part B

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1 Composites: Part B 42 (2011) Contents lists available at ScienceDirect Composites: Part B journal homepage: Low-grade glued laminated timber reinforced with FRP plate Gary M. Raftery, Annette M. Harte Civil Engineering, College of Engineering and Informatics, National University of Ireland, Galway, Ireland article info abstract Article history: Received 8 April 2010 Received in revised form 19 January 2011 Accepted 25 January 2011 Available online 1 February 2011 Keywords: A. Wood A. Glass fibres A. Polymer matrix composites (PMCs) D. Mechanical testing Fibre-reinforced polymers (FRPs) are an accepted material by structural engineers for the strengthening of structural elements. With increasing environmental concerns, further emphasis is being placed on the use of sustainable construction materials such as timber. This paper describes an experimental programme which examines the reinforcement in flexure of low-grade glued laminated timber (glulam) with a commercially attractive recyclable FRP. Unreinforced glulam beams, FRP plate reinforced glulam beams, unreinforced glulam beams incorporating an additional sacrificial lamination and FRP plate reinforced glulam beams, which had a sacrificial lamination bonded beneath the FRP plate, were fabricated and tested in flexure. The unreinforced glulam demonstrated linear elastic behaviour and exhibited brittle tensile-flexural failures in comparison to the pseudo-ductile behaviour of the reinforced beams. The addition of reasonable reinforcement percentages strategically located in the tension zone resulted in moderate enhancements in the stiffness while more significant improvements in the ultimate moment capacity were obtained. Strain profiles illustrated that increased utilisation of the compression strength of the timber is obtained when FRP reinforcement is included. Ó 2011 Elsevier Ltd. All rights reserved. 1. Introduction In the past two decades, structural engineers have become increasingly aware of the advantages that are offered by the use of advanced fibre reinforced polymer (FRP) materials. The strength and stiffness of these materials are significantly high in relation to their weight when compared to other construction materials and have obtained favourable results in a large number of studies in the reinforcement or structural rehabilitation of concrete and steel members with particular attention directed to the former. FRPs materials also portray superior durability over conventional construction materials [1,2]. As a result maintenance costs are reduced with their use. Furthermore, they portray excellent resistance to fatigue. To date, limited studies have been carried out on the reinforcement of structural timber as the primary element. FRP materials are more suitable for the reinforcement of timber in comparison to steel because of the susceptibility of steel to corrosion when in contact with the moisture content of the timber. As one of the oldest building materials, timber possesses many applications in the construction industry. It holds many advantages as a structural material [3,4]. It is a natural renewable resource with a secure supply, recyclable, cost competitive, strong and is aesthetically pleasing. With increasing efforts to promote the use of sustainable materials such as timber in the construction industry, considerable potential is associated with the bonding of FRP plate Corresponding author. Tel.: ; fax: address: Annette.Harte@nuigalway.ie (A.M. Harte). to low-grade timber. The technique can be easily and effectively integrated into the glued laminated timber (glulam) manufacturing procedure and adds negligible depth and mass to the member that is being reinforced. Glass fibre reinforced plastics (GFRP) which incorporate E-glass is the most appropriate type of fibre reinforcement because of its low cost to good mechanical properties in comparison to the use of other glass fibre types, aramid fibres and carbon fibres FRP reinforced wood A number of research studies have examined the attractive option of reinforcing wooden flexural members with fibre-reinforced polymer plates. Although this idea has existed for some time the amount of research undertaken in this area remains limited. Theakston [5] commented on the possible feasibility and practicality of FRP reinforced timber. Spaun [6] showed that fibreglass could be used to substantially stiffen and increase the tensile strength of finger-jointed western hemlock cores. Bulleit [7] compiled a literature review on several investigations into the reinforcement of timber but the majority of these techniques have not achieved commercial viability because of the cost resulting from the additional manufacturing steps involved. Van de Kuilen [8] reported that the stiffness can be considerably increased by using glass fibre reinforced polyester profiles. Furthermore, a considerable increase in ultimate bending moment was achieved although problems were reported in the bond quality between the FRP material and the timber. Moulin et al. [9] discovered that by incorporating /$ - see front matter Ó 2011 Elsevier Ltd. All rights reserved. doi: /j.compositesb

2 G.M. Raftery, A.M. Harte / Composites: Part B 42 (2011) fibreglass in a glulam beam, increases in the flexural rigidity and loading capacity could be achieved. Furthermore, it was observed that the reinforced beam can sustain very large strains, while maintaining its load carrying capacity after rupture of the outer lamination. Tingley and Cegelka [10] discussed numerous material considerations that arise with the reinforcement of wood when using FRP materials but concluded that substantial strength improvement and economic benefits can result with the technique. They further remarked that careful selection of the correct FRP material type for end use is important. Furthermore, Tingley [11] stated that the most important benefits from FRP reinforcement in the tension zone were strength enhancement, increased MOE, the use of less wood fibre for specific load carrying capacity, the use of low-grade wood in place of high-grade wood and a reduction in variability. Dagher et al. [12] reported from a test programme which examined the bonding of FRP plates between a sacrificial strip and outer tensile wood lamination that the largest increases in strength were obtained with lower grades of wood and therefore the highest value added benefits would result in these circumstances. In another study it was seen that by reinforcing glulam with a material with higher stiffness and strength such as fibre-reinforced polymers, a reduction could be made in the depth while achieving an increase in the bending strength and stiffness [13]. Carbon fibre plate reinforcement was used to reinforce glulam beams where the expected mechanical improvements were obtained but it was believed that the gains in strength over that of unreinforced beams would not be economically viable because of the cost of carbon fibre reinforcement [14]. However, there was much less variability in the performance of the reinforced beams compared to the unreinforced beams. Johns and Lacroix [15] reported that wood adjacent to the composite material experienced an increase in effective strength in comparison to when no composite material was present. This behaviour was thought to be due to the composite material acting as a bridge across local defects. Gilfillan et al. [16] showed that for modest carbon fibre plate reinforcement percentages, a considerable increase in both bending stiffness and ultimate load capacity was achieved. However, it was further concluded from the test programme undertaken that it would not be economical in terms of stiffness enhancements obtained to reinforce with carbon fibre plate reinforcement. Fiorelli and Alves Dias [17] found that beams reinforced with carbon fibres and glass fibres resulted in a significant gain in ductility in comparison with non-reinforced beams. Furthermore, the reinforcement gave greater reliability to the beam, and it was reported that there was a reduction in tensile failures initiated by defects. External bonding of carbon fibre sheets to reinforce old wood beams was carried out by Borri et al. [18] who reported increases in flexural stiffness and capacity. Both steel and carbon fibre reinforced plates were investigated by Issa and Kmeid [19]. The reinforced beams showed a more ductile mode of failure and an increased load carrying capacity. The advantages of the lightweight CFRP with its ease of handling in comparison to steel were highlighted. An economic analysis executed by Stevens and Criner [20] concluded that savings would be greater for deep beams that required high bending strength. There have been some successful efforts to commercialise FRP-reinforced glulam in both North America and in Europe [21]. Research work has also been carried out on the reinforcement of wood using internal rod and bar FRP reinforcement. Gentille et al. [22] stated that the failure mode changed from brittle tension to compression failure with the use of near surface GFRP bars with an increase in flexural strength of between 18% and 46% as the reinforcement overcame the effect of local defects in the timber. Reduced variability and increased ultimate strength were recorded by Svecova and Eden [23] with the application of near surface mounted GFRP bars in timber stringers. Alam [24] stated that lower cost, low modulus bonded-in reinforcements which incorporated glass fibres had as much potential for improving the flexural properties of timber as higher modulus reinforcements such as steel or FRPs which included carbon fibres. In another study which included CFRP rods it was seen that experimental results showed a significant increase in ultimate strength and stiffness [25]. Despite favourable results obtained with bonded in reinforcements, it is considered that this reinforcing technique would be more suitable for repair and upgrade applications because of the routing that is involved rather than fabrication of a new commercial product Objectives of the present study The objective of this study is to examine the reinforcement of low-grade glulam using a pultruded glass fibre reinforced polymer composite plate. The reinforcement procedure is designed to take place in a commercial workshop environment during the glulam manufacturing. The experimental test programme involved the fabrication and testing in flexure of unreinforced beams and reinforced beams both with and without the addition of a sacrificial lamination. The mechanical performance of the reinforced beams is compared with that of the unreinforced glulam beams with regard to the load deflection behaviour, failure mode, enhancements in both stiffness and ultimate moment capacity as well as the changes in strain profile distribution. 2. Materials 2.1. Timber The timber species used was Irish grown Sitka spruce. The timber stock was C16 grade in accordance with EN 338 [26], plain sawn cut and was kiln dried in the sawmills to approximately 18% moisture content. All the timber was harvested from the same stand in order to reduce variability from contrasting environmental conditions during growth. The boards were 4200 mm in length with a nominal section size of 96 mm 44 mm. The boards were stored for three months in a conditioned environment of 65 ± 5% relative humidity and 20 ± 2 C temperature once delivered to the laboratory. An approximate equilibrium moisture content of 12% was obtained after the conditioning period and a mean board density of 390 kg/m 3 was recorded Fibre reinforced polymer The FRP material used in the test programme comprised glass fibres in an engineered thermoplastic polyurethane matrix [27]. Selection was made based on its mechanical properties, availability and commercial viability for the reinforcement of the low-grade timber. This material was manufactured using the pultrusion process and incorporated glass fibres aligned unidirectionally. The mechanical properties of the composite plate in comparison to Irish grown Sitka spruce are shown in Table 1. Good potential is offered for the reinforcement of low-grade glulam if the reinforcement is strategically located in the higher stressed tensile zone of the member. When subject to a normal internal service environment, it is envisaged that the ductile matrix would facilitate better load sharing between the fibres than if a brittle thermosetting resin was used. Thereby, creep rupture life is improved. The FRP plate used in the testing involved the epoxy bonding together of two 96 mm wide, 1.2 mm thick plates to form an overall plate thickness of 2.8 mm. FRP reinforcement percentages of approximately 1.26% and 1.12% were correspondingly applied to the 190 mm deep and 215 mm deep glulam beams.

3 726 G.M. Raftery, A.M. Harte / Composites: Part B 42 (2011) Table 1 Mechanical properties of FRP and timber. Material Ultimate tensile strength (N/mm 2 ) Modulus of elasticity in tension (N/mm 2 ) Ultimate compressive strength (N/mm 2 ) FRP 1000 a 45,000 a 701 b 38,440 b Irish grown Sitka spruce 23.7 c 8111 c d 8000 e a FRP material properties as reported by manufacturer [27]. b FRP material properties as tested for two 1.2 mm plates bonded using a 0.4 mm epoxy bondline [28]. c In-grade testing with mean moisture content of 12% and mean density of 403 kg/m 3 [28]. d In-grade testing with mean moisture content of 11.9% and mean density of 405 kg/m 3 [28]. e In-grade testing with mean moisture content of 12% and mean density of 449 kg/m 3 [16]. Modulus of elasticity in compression (N/mm 2 ) 2.3. Adhesives The adhesives that were selected for the bonds between the timber laminations and at the FRP-wood bond interface were examined in previous research programmes carried out by Raftery et al. [29 31]. The research on the FRP-wood bond interface was undertaken using a methodical approach whereby shear strength performance and adherend percentage failures were contrasted for solid control specimens, wood wood bonded specimens, nonmoisture cycled FRP-wood bonded specimens and moisture cycled FRP-wood bonded specimens. The bond quality of six conventional wood laminating adhesives and three structural epoxy adhesives was studied in association with a silane based adhesion promotor. Performance was assessed using a compressive lap shear specimen. From the results obtained from this research, a phenol resorcinol formaldehyde (PRF 1) was selected for the bonding of the wood-wood laminations. The epoxy, a well-recognised civil engineering two-component thixotropic adhesive, was the adhesive which showed best performance when bonding to the FRP material that was to be used as the reinforcement for the glulam beams in this study [32]. 3. Experimental methods 3.1. Test programme Table 2 Test programme. Phase Arrangement Depth (mm) Repetitions Beam numbers A Unreinforced , 2, 3, 36, 37, 38 B Reinforced , 11, 18, 24, 31 C Unreinforced , 9, 10, 16, 17, 22, 23, 29, 30, 33 D Reinforced , 15, 21, 28, 35 Fig. 1. Beam configurations: (a) Phase A. (b) Phase B. (c) Phase C. (d) Phase D. The experimental programme involved the fabrication and testing for both stiffness and ultimate moment capacity of unreinforced and reinforced glulam beams. The test programme is shown in Table 2 and the configurations are shown in Fig. 1. The 215 mm deep beams (Phases C and D) involved the bonding of a sacrificial lamination, so that charring resistance was provided by the timber to the FRP plate in case of exposure to fire. The thickness selected was in excess of the 20 mm which was required to give 30mins resistance to charring as stated in BS5268 [33]. Martin and Tingley reported that fire performance can be improved by placing the FRP internally as the wood insulates the FRP thus delaying the polymer matrix reaching its glass transition temperature [34]. Williamson showed by experimental testing that a one hour fire resistance rating can be obtained by larger size FRP-reinforced glulam beams even when no sacrificial lamination is included [35]. However, the section size employed in the present test programme was considerably smaller than that used by Williamson. A larger number of unreinforced beams were fabricated because of the greater variability that would be associated with their behaviour. Ten Phase C beams were fabricated in comparison to six Phase A beams as it was envisaged that Phase D beams were the finished product and consequently it was deemed more important to compare the performance of a greater number of unreinforced beams of the same depth. The beam test programme involved selectively staggering the quality of the unreinforced beams and reinforced beams so that a more direct comparison of the behaviour of the beams could be obtained. By following this procedure, which is explained in greater detail in Section 3.2, the effect of reinforcing both good quality and poor quality beams could be assessed. Furthermore, a more direct comparison of the reinforced beams was made with the unreinforced beams by comparing with the mechanical performance of both good quality and poor quality beams.

4 G.M. Raftery, A.M. Harte / Composites: Part B 42 (2011) Beam fabrication The timber stock was initially assessed by mechanical grading [36]. Subsequently, visual grading of the boards was undertaken with reference to IS 127 [37]. Boards associated with excessive distortion, which were thought to be unsuitable for use as laminations, were removed from the lamination stock. Boards which had excessive margin knot area ratios, excessive total knot area ratios or other critical strength reducing defects such as fissures were not used in the extreme tension and compression laminations. Boards were ranked on the basis of the readings obtained from the mechanical grader after taking into consideration the magnitude of the bending moment to which they would be subjected along their length during the four point flexure testing. The boards were knife planed to 38 mm thick laminations with each beam incorporating five laminations and having an initial total beam depth of 190 mm. As the beams were to be tested in four point bending with a span-depth ratio of 18:1, this gave a span of 3420 mm [38]. The laminations were trimmed from one or both ends to the required length such that the quality of the material in each beam that would be subjected to the zone of maximum bending moment would be optimised. The lay-up for all the glulam beams involved using the best quality laminations as the most extremely stressed tension laminations at the bottom of each beam. The ranking procedure involved locating the best quality lamination from the stock at the soffit of Beam 1 with the next best quality lamination being placed at the bottom of Beam 2. This consecutive sequence was maintained as the beam number ascended. An identical procedure was used for the next set of laminations which were used in the critical compressive zone at the top of the beams. Subsequently, the secondary tension laminations were allocated followed by the secondary compressive laminations with both sets of laminations using the ranking procedure described above. The remaining laminations which were considered the weakest based on the results from the grading process were placed in the core of the beam at the position of the neutral axis where stresses would be at a minimum. Based on this manufacturing methodology, higher quality laminations were associated with lower beam numbers. Consequently, by following this ordering system for the lay-ups, beams which were manufactured at the start of the programme were theoretically stronger and stiffer than the beams manufactured later in the programme. No more than two hours was permitted to elapse between the planing process and application of the PRF adhesive at a spread rate of 400 g/m 2 as studied previously [30,31]. The assembly was subsequently left clamped at a pressure of 0.7 N/mm 2 for 24 h in a clamping jig in an environment of 65 ± 5% relative humidity and 20 ± 2 C temperature. In all the manufactured beams, the laminations were orientated such that the pith was towards the top side of the board and followed the general orientation as specified in EN 386 [39]. The bonding of the FRP reinforcing plates in Phases B and D involved initially knife planing the soffit of the glulam beam as for the bonding of the timber laminations. The surface preparation of the FRP involved gentle abrasion using 320 grade wet and dry emery paper so that the top layer of the material, which was likely to contain release agents used in the manufacturing process and other contaminants were removed. The material was subsequently wiped clean with methylated spirits before the epoxy adhesive was applied. A bondline thickness of 0.5 mm was used and the arrangement was clamped as stated before. A durable high quality bond was achieved to this material using these techniques in a previous study [31]. The series of sacrificial laminations, which were used in the 215 mm deep unreinforced beams (Phase C) and which were bonded below the FRP plates in the 215 mm reinforced beams (Phase D), were similarly graded both mechanically and visually and the best quality laminations were assigned with the lower numbers. Bond preparation involved using a 0.5 mm epoxy bondline thickness and the surface preparations of the timber and FRP plate as previously stated. The grain orientation of the sacrificial lamination was opposite to that of the beam in order to compensate for cupping in the timber. Because of the varying density of the laminations, minor variations occurred in the thickness during the planing process and hence the laminate thickness and beam depth was recorded for each beam in the four phases examined using a vernier calliper with a precision of 0.01 mm Beam testing All the beams (Phases A D) were initially tested in four point bending for both local and global stiffness in the 190 mm deep unreinforced state in accordance with EN 408 [38]. The test setup is as shown in Fig. 2. Testing was carried out using a Dartec 500 kn machine when conditions of 65 ± 5% relative humidity and 20 ± 2 C temperature were present. The local stiffness is determined using a linear variable differential transformer (LVDT) positioned in a hanger suspended from the neutral axis over a distance of five times the depth of the beam (Fig. 2), which measures the deflection in the zone of maximum bending moment. The measurement of local stiffness is not affected by shear deflection or indentations in the timber at the supports. The global stiffness measurement is recorded from a LVDT inverted on the compression face of the beam, which measures deflection over the entire length of the member relative to the supports. The beams were loaded at a constant crosshead displacement rate of 0.57 mm/s after a preload of 300 N was applied prior to the test [38]. During the test, it was ensured that the load did not exceed 40% of the ultimate load or exceed the elastic limit. Lateral supports were positioned approximately 300 mm outside of the loading heads and frictional effects were reduced to a minimum by the use of polytetrafluoroethylene (PTFE) strips sliding over each other. Steel plates of dimensions 95 mm 90 mm 10 mm were placed beneath the two loading heads and at the supports so that local indentations were minimised. The test arrangement, shown in Fig. 2, was also employed for testing of the beams reinforced with the FRP plate. However, the hanger to accommodate the LVDT for the measurement of the local stiffness was at a lower depth in the section of the beam because of the shift in neutral axis. Fig. 2. Stiffness test arrangement for Phase A and Phase B beams (arrangement for Phase C and Phase D beams shown in brackets).

5 728 G.M. Raftery, A.M. Harte / Composites: Part B 42 (2011) The test arrangement for the 215 mm deep beams in Phases C and D, which include the sacrificial lamination, is also illustrated in Fig. 2. The position of the loading heads on the Dartec machine was adjusted to correspond to the test standard and each beam was tested at a constant crosshead displacement rate of mm/s. Again it was ensured that the load did not exceed 40% of the ultimate load or go above the elastic limit. The hanger for the neutral axis was at a lower position for the reinforced beams because of the neutral axis being at a lower position. Both local and global stiffness testing was undertaken with two LVDTs as shown in Fig. 3. Strength testing of all the beams was executed after unloading of the beams and removal of the local stiffness LVDT and hanger. Fracture of the specimens was initiated in 300 ± 120 s [38]. The strain distribution behaviour was recorded on two of the 215 mm deep beams, one of the Phase C beams and one of the Phase D beams, using the configuration for the strain gauges as illustrated in Fig. 4. Ten strain gauges of 60 mm gauge length (PL-60-11) were placed at midspan throughout the depth of each beam, on both faces of the timber laminations. By applying strain gauges to both sides of the beam, twist that occurred during flexure could be accounted for. Gauges of 60 mm in length allowed a mean estimation of the stresses to be obtained rather than a spot value and were carefully applied to clear wood on surface of timber so that defects would not disturb the readings being taken. Fig. 5 shows Beam 17 with strain gauges attached on one face prior to testing. 4. Results and discussion 4.1. Load deflection behaviour Load deflection behaviour, Phase A Unreinforced 190 mm deep beams The load deflection behaviour to failure for the 190 mm deep unreinforced (Phase A) is shown in Fig. 6. The beams in Phase A exhibited linear elastic behaviour to failure. Because of the low grade of the timber used in the study, defects such as knots were still present in the zone of maximum bending moment in both the tension and compression laminations despite the grading process used during manufacturing of the beams. Timber with defects has a lower tensile strength than compressive strength [26]. Therefore, even when using better quality laminations at the bottom of the beam than at the top of the beam, the yield stress in tension is generally exceeded before the yield stress in compression is reached. As a result nonlinear behaviour in the compression zone is not experienced in these beams. All the beams demonstrated brittle failures which resulted from excessive tensile stresses in the bottom lamination. With the exception of Beam 37 which failed in clear wood, all the failures were initiated from knots at the soffit of the beam and the beams were severely damaged after the testing procedure. The catastrophic nature of the failure of the unreinforced beams is demonstrated in Figs. 7a and b which illustrate Beam 1 during test execution, and after failure. The failed specimen is severed through the midspan of the member in the testing machine. The presence of shear cracks after the initial tensile fractures was noted in a number of the beams. Fig. 8 illustrates a tensile failure with shear cracking in Beam 3. No compression buckling was evident in the top lamination of any of the beams. Fig. 3. LVDTs for local and global stiffness testing Load deflection behaviour, Phase B Reinforced 190 mm deep beams The 190 mm deep beams that were reinforced with the FRP plate (Phase B) demonstrated pseudo-ductile behaviour, particularly in Beam 5 (Fig. 9). Failure in Beam 5 occurred in the clear wood in the mostly highly stressed tension lamination. Beams 11, 18 and 24 failed at knots in the tension laminations. Limited ductility was shown by Beam 31 as this beam failed at a significantly large knot in the bottom lamination. Therefore, the amount of ductile behaviour in the FRP plate reinforced beams depended Fig. 4. Strain gauges arrangement (a) Phase C and (b) Phase D.

6 G.M. Raftery, A.M. Harte / Composites: Part B 42 (2011) Fig. 7b. Phase A. Failed specimen (Beam 1). Fig. 5. Strain gauges arrangement (Beam 17). Fig. 8. Phase A. Tensile failure with associated shear fractures (Beam 3). Fig. 6. Load deflection behaviour for unreinforced 190 mm deep beams (Phase A). Fig. 9. Load deflection behaviour for FRP plate reinforced 190 mm deep beams (Phase B). Fig. 7a. Phase A. Test execution. on the quality of the bottom timber laminations as when a better quality lamination is placed adjacent to the FRP plate in the critical tension zone, the yield stress in compression may be reached before the yield stress in tension is exceeded. This is because the FRP plate contributes to the tensile capacity of the beam and helps to dissipate stresses away from the critical strength reducing defects in the adjacent bottom timber lamination Load deflection behaviour, Phase C Unreinforced 215 mm deep beams The load deflection behaviour of the beams in Phase C is shown in Fig. 10. All of the unreinforced 215 mm deep beams failed in tension in the bottom lamination. Eight of the ten beams failed at knots, which were located in the zone of maximum moment between the loading head in the bottom 25 mm lamination and all of these beams demonstrated linear elastic behaviour to ultimate moment capacity. After these initial fractures, shear cracks were evident and in some cases propagated along the annular ring, past the loading heads and towards the supports as seen in Fig. 11 for Beam 23. Two beams failed in clear wood in the bottom lamination. One of these beams, Beam 29, demonstrated limited ductile behaviour. Beams 9 and 10 were of a greater depth than the other

7 730 G.M. Raftery, A.M. Harte / Composites: Part B 42 (2011) Fig. 10. Load deflection behaviour for unreinforced 215 mm deep beams incorporating sacrificial lamination (Phase C). to the ultimate moment capacity of the beams. The remaining two beams, Beam 21 and Beam 35, did not recover to exceed the load at which the sacrificial lamination fractured. Compression buckling was clearly visible at a knot in the top lamination of Beam 35 as shown in Fig. 13. This physical behaviour, as well as the pseudoductility visible in the load deflection plots demonstrates that by strategically positioning FRP plate of sufficient, but reasonable, reinforcement percentage in the more highly stressed tensile region of the beam, pseudo-yielding of the compression fibres in the timber at the top lamination of the beam can occur. This behaviour permits for a more efficient use of structural material in the compression laminations before tensile rupture occurs either at a knot or in clear wood at the bottom of the structural member. Upon fracturing of the sacrificial lamination in Beam 35, the lamination debonded as shown in Fig. 14 and the load deflection behaviour of the member was similar to that of the 190 mm deep reinforced beams and did not recover load greater than that at which the fracture occurred Stiffness Stiffness testing, Phase A Unreinforced 190 mm deep beams The results for both global and local stiffnesses of the Phase A beams is shown in Fig. 15. The beams had a mean global stiffness of Nmm 2 with standard deviation of Nmm 2 and a mean local stiffness of Nmm 2 with standard deviation of Nmm 2. The average depth of the Fig. 11. Phase C. Tensile failure with associated shear cracks propagating along the grain (Beam 23). beams in Phase C and consequently were associated with a greater stiffness as discussed in Section Load deflection behaviour, Phase D Reinforced 215 mm deep beams The load deflection curves of the reinforced 215 mm deep beams, which included the sacrificial lamination (Phase D) showed pseudo-ductile beam behaviour as can be seen in Fig. 12. All the beams experienced fracturing of the sacrificial lamination as a result of the excessive tensile stresses experienced. For three of the five beams (Beam 8, Beam 15, Beam 28), the ultimate bending moment capacity was obtained after the beams initially fractured in the sacrificial lamination. Considerable ductile behaviour was experienced by these beams after this initial fracture. The fractures in the sacrificial lamination of these beams did not cause delamination and it is believed that this laminate continued to contribute Fig. 13. Phase D. Compression buckling at a knot in the top lamination (Beam 35). Fig. 12. Load vs. deflection behaviour for FRP plate reinforced beams incorporating 25 mm sacrificial lamination (Phase D). Fig. 14. Phase D. Fracturing of sacrificial lamination (Beam 35).

8 G.M. Raftery, A.M. Harte / Composites: Part B 42 (2011) Fig. 15. Stiffness testing. Phase A: Beams 1, 2, 3, 36, 37 and 38. Key: A = Global stiffness for 190 mm unreinforced beams and B = Local stiffness for 190 mm unreinforced beams. beams in Phase A was mm with standard deviation of 1.87 mm. Beam 38 was marginally deeper ( mm) than the other beams in this category and hence was associated with a greater stiffness. Differences in the global and local stiffness results were believed to be primarily as a result of the presence of shear deformation in the global stiffness which does not arise in the local stiffness test. Reduced stiffness may also be associated with the global stiffness results because of the likelihood of indentation at the supports resulting from the low density of the softwood used to manufacture the glulam beams. Furthermore, from the grading procedure undertaken, better quality material was located in the zone between the loading heads where the hanger for the LVDT to measure the local stiffness is located Stiffness testing, Phase B Reinforced 190 mm deep beams The stiffnesses of the beams in Phase B, the 190 mm deep beams, which were tested both in the unreinforced state and tested again when reinforced with the FRP plate, are shown in Fig. 16. In the unreinforced state, the beams had a mean global stiffness of Nmm 2 with standard deviation of Nmm 2 and a mean local stiffness of Nmm 2 with standard deviation of Nmm 2. Both these results are in the same range as Phase A, the unreinforced 190 mm deep beams, which demonstrates how the beams were methodically manufactured. The mean depth of the beams in this state was mm with standard deviation of 0.94 mm. Fig. 16. Stiffness testing. Phase B: Beams 5, 11, 18, 24, and 31. Key: A = Global stiffness for 190 mm deep unreinforced beams; B = Local stiffness for 190 mm deep unreinforced beams; C = Global stiffness for 190 mm deep reinforced beams; and D = Local stiffness for 190 mm deep reinforced beams. After the reinforcement was added in the tension zone, the mean global stiffness of the beams increased to Nmm 2 with standard deviation of Nmm 2, which represents a mean increase of 12.13% with standard deviation of 4.91% in comparison to the unreinforced beams. It can be seen in Fig. 16 that there is a consistent improvement for all the beams tested. This improvement occurs because of the high tensile modulus of elasticity associated with the FRP plate (Table 1). By strategically locating the plate at an optimum distance away from the neutral axis in the zone of higher stressed tensile wood fibres, an improvement is obtained for the stiffness of the hybrid beam. The mean local stiffness of these beams when reinforced was Nmm 2 with a standard deviation of Nmm 2. This represents a percentage increase of 13.15% with standard deviation of 3.68%. The local stiffness results are of a higher magnitude than the global stiffness results as previously explained for the Phase A beams except that no indentation occurs at the supports because the FRP plate is positioned on the supports. The depth of the reinforced beams was mm with standard deviation of 0.22 mm. This represents a difference in depth of 0.47% between the reinforced beams (Phase B) and unreinforced beams (Phase A). If a solid plate 2.8 mm in thickness was used as the reinforcement rather than the two 1.2 mm thick plates that were bonded together using a 0.4 mm epoxy bondline which were used in this test programme, further increases in stiffness would be obtained for the same depth beams Stiffness testing, Phase C Unreinforced 215 mm deep beams The results of the global and local stiffness testing of the beams in Phase C can be seen in Fig. 17. The mean global and local stiffnesses of the beams before the sacrificial laminations were added were Nmm 2 and Nmm 2 with standard deviations of Nmm 2 and Nmm 2 respectively. The mean global and local stiffnesses of the beams when the sacrificial laminations were included were Nmm 2 and Nmm 2 with standard deviations of Nmm 2 and Nmm 2, respectively. Mean beam depth without the sacrificial lamination was mm with standard deviation of 2.95 mm in comparison to mm and standard deviation of 2.52 mm when the sacrificial laminations were added. It can be seen that Beam 9 and Beam 10 were associated with a significantly higher stiffness than the other beams in the category because together they have a mean depth of mm, which was 3.93 mm greater than mm. The stiffer sacrificial laminations, which were also associated with these beams because of the manufacturing procedure previously explained, would also have a contribution to the superior stiffness of these two beams. Variations between the global and local stiffness measurements are as a result of the reasons as stated for the beams in Phase A Stiffness testing, Phase D Reinforced 215 mm deep beams The mean global stiffness of the Phase D beams when no FRP reinforcement was bonded to the tension zone was Nmm 2 with standard deviation of Nmm 2. The corresponding mean local stiffness was Nmm 2 with standard deviation of Nmm 2. These results demonstrate that the quality of the laminations used in the beams in Phase D were in the same range as used in Phase A, Phase B and Phase C. The mean depth of the beams when unreinforced was mm with a standard deviation of 1.46 mm. When the reinforcement was bonded onto the bottom of the beams a mean increase of 10.1% and 9.8% was obtained in the global and local stiffness respectively as can be seen in Fig. 18. Standard deviations of 0.96% and 2.36% were associated with these results. These percentage increases are consistent with that obtained for the beams in Phase B. The mean depth of the beams when reinforced

9 732 G.M. Raftery, A.M. Harte / Composites: Part B 42 (2011) Fig. 17. Stiffness testing. Phase C: Beams 4, 9, 10, 16, 17, 22, 23, 29, 30 and 33. Key: A = Global stiffness for 190 mm deep unreinforced beams; B = Local stiffness for 190 mm deep unreinforced beams; E = Global stiffness for 215 mm deep unreinforced beams; F = Local stiffness for 215 mm deep unreinforced beams. Fig. 18. Stiffness testing. Phase D: Beams 8, 15, 21, 28 and 35. Key: A = Global stiffness for 190 mm deep unreinforced beams; B = Local stiffness for 190 mm deep unreinforced beams; C = Global stiffness for 190 mm deep reinforced beams; D = Local stiffness for 190 mm deep reinforced beams; G = Global stiffness for 215 mm deep reinforced beams; and H = Local stiffness for 215 mm deep reinforced beams. with the plate was mm with a standard deviation of 1.06 mm. When the sacrificial lamination was added below the FRP plate, the mean global stiffness was increased by 55% to Nmm 2 with standard deviation of Nmm 2. The mean local stiffness was increased by 47% to Nmm 2 with standard deviation of Nmm 2. The beams in this state had a mean depth of mm which represents a percentage difference in depth of 1.1% between the reinforced beams in Phase D and the unreinforced beams in Phase C. The standard deviation associated with the depth was 0.98 mm. Variations between the global and local stiffness measurements are due to the presence or absence of shear deflection, local indentation at the supports and the procedure undertaken for the fabrication of the beams where the better quality material was located in the zone of maximum bending moment Ultimate moment capacity Table 3 Ultimate bending moment, M ult : 190 mm deep beams. Phase A Beams M ult (knm) Phase B Beams M ult (knm) Mean Mean St. Dev St. Dev The ultimate moment capacity M ult achieved by each of the 190 mm deep beams is shown in Table 3. It can be seen that the mean ultimate capacity of knm for the Phase B reinforced beams is considerably greater than that of knm for the Phase A unreinforced beams, which represents an increase of 38%. Furthermore, a lower standard deviation of 3.55 knm is associated with the reinforced beams in comparison to 3.92 knm for the unreinforced beams. This suggests that a reduction in the variability of the ultimate moment capacity of the beam occurs by bonding the FRP reinforcement plate onto the soffit of the beam. The results of the failure tests for the 215 mm deep beams in terms of the ultimate moment capacity, M ult, are shown in Table 4. It is seen that a mean increase in moment capacity of 28.6% is obtained by the Phase D beams, which incorporate the FRP reinforcement in comparison to the Phase C unreinforced beams. Furthermore, a lower standard deviation of 2.86 knm is associated with the reinforced beams in comparison to 3.6 knm demonstrating the reduction in variability of ultimate moment capacity of the reinforced beams.

10 G.M. Raftery, A.M. Harte / Composites: Part B 42 (2011) Table 4 Ultimate bending moment, M ult : 215 mm deep beams. Phase C Beams M ult (knm) Phase D Beams M ult (knm) Mean Mean St. Dev St. Dev Strain profile distribution The strain profile for Phase C in which Beam 17 is instrumented is shown in Fig. 19. Phase C comprised the unreinforced 215 mm deep beams. It is believed that Beam 17 was a good representation of the mechanical properties of the beams in Phase C as the laminations used to fabricate this beam are close to average of the mechanical properties of the laminations used to fabricate all of the unreinforced 215 mm deep beams. It can be seen that the strain behaviour is linear to failure. The maximum compressive strain of recorded by Strain Gauge 1 at failure is low in comparison to maximum tensile strain of recorded by Strain Gauge 10. Differences in strain readings between gauges on the same lamination but on different faces are not significant. Fig. 20 illustrates the damaged gauges on Beam 17 after failure. The strain profile for the Phase D beam, Beam 21, which was reinforced with FRP plate and included a sacrificial lamination is shown in Fig. 21. It can be seen that much greater strain is recorded in the top compressive lamination of this beam in comparison to that experienced by the unreinforced glulam. This is illustrated by the readings from Strain Gauge 1. At an applied moment of 20 knm, a strain reading of was obtained in the beam with the FRP plate included in comparison to a strain of for the unreinforced beam. This illustrates that by strategically including the FRP reinforcement in the more highly stressed tension region at the bottom of the beam, better utilisation of the compressive properties of the timber is obtained. It should be noted that Strain Gauge 1 did not record the maximum strain at the top of the FRP plate reinforced beam as this plasticization occurred at a localised defect away from the midspan. Nonlinear compression behaviour is recorded by the strain gauges positioned above the neutral axis when the moment is greater than 25 knm. No such strain behaviour was seen for the gauges on the unreinforced beam. The position of the neutral axis does not significantly deepen in the beam as the applied moment is increased and plasticization of the top compression lamination occurs. Strain Gauge 10, which was positioned at the bottom of the beam, so that it could record the maximum tensile strain, shows nonlinear behaviour but this is believed to be as a result of micro-fractures occurring in wood fibres of the sacrificial lamination before ultimate load capacity of the beam was obtained. The strain profile indicates that there is no appreciable slip between the timber and reinforcing FRP material. Deviations in strain readings between the two faces of the beams are not considered to be significant. 5. Costing Fig. 20. Phase C. Strain gauge arrangement (Beam 17). Xu carried out a cost analysis on FRP-glulam bridge panels based on tender prices and found that the GFRP material cost was about 24% of the unreinforced glulam cost for each 1% of reinforcement used [40]. Meier et al. reported GFRP material costs of 39% of the unreinforced glulam for each 1% of reinforcement used. Based on the higher of these two costs, the GFRP costs for the 190 mm and 215 mm beams in the current study are 49% and 44% of the glulam costs [41]. The corresponding increases in moment capacity are 38% and 29%, respectively. 6. Theoretical analysis Fig. 19. Phase C. Strain profile (Beam 17). A transformed section analysis was undertaken to predict the flexural stiffness and ultimate moment capacity in which a composite elastic section was assumed. The mechanical properties of the reinforcement and the mean mechanical properties of the timber that were determined from mechanical testing [28] and given in Table 1 were used in the analysis. The laminating effect which is the increase in apparent strength of a timber lamination when bonded in a glulam beam in comparison to when tested in a uniaxial test was accounted for in the model. Falk et al. [42] reported a value of 1.26 for the laminating effect for a beam made of homogeneous C30 grade laminations. Use of higher grade C37 laminations

11 734 G.M. Raftery, A.M. Harte / Composites: Part B 42 (2011) Fig. 21. Phase D. Strain profile (Beam 21). Table 5 Experimentally determined stiffness versus theoretically predicted stiffness values. Beam Phase Experimental EI (N mm 2 ) Theoretical EI (N mm 2 ) Experimental EI/theoretical EI Phase A 4.86E+11 (11E+10) 4.52E+11 (1.33E+10) 1.08 Phase B 5.67E+11 (1.9E+10) 5.30E+11 (1.93E+09) 1.07 Phase C 7.06E+11 (7.5E+10) 6.26E+11 (2.24E+10) 1.13 Phase D 7.73E+11 (3.3E+10) 7.04E+11 (7.53E+09) 1.10 Table 6 Experimentally determined ultimate moment capacity versus theoretically predicted ultimate moment capacity. Beam Phase Experimental M ult (knm) Theoretical M ult (knm) Experimental M ult /theoretical M ult Phase A (3.92) (0.34) 1.24 Phase B (3.55) (0.05) 1.35 Phase C (3.61) (0.52) 1.18 Phase D (2.86) (0.18) 1.30 resulted in a significantly lower laminating factors [43]. The application of 1.26 to the tensile strength of the C16 grade laminations is conservative. The results for the flexural stiffness and ultimate moment capacity are given in Tables 5 and 6, respectively. Strong agreement was obtained between the predicted flexural stiffness and experimentally determined flexural behaviour. All predictions were conservative and this trend is believed to be as a result of the dispersion and stress redistribution effects present in glued laminated timber [43]. Further improvement could be achieved if the modulus of elasticity of each lamination was considered in the analysis rather than mean values determined from axial test programmes. The predictions for the ultimate moment capacity of all the beams in all four beam phases were also conservative. For the unreinforced beams, this is believed to be as a result of the conservative lamination factor that was employed in the model. Furthermore, the model uses a mean tensile strength for the timber whereas in reality the bottom lamination in of higher quality because of the manufacturing procedure undertaken. Predictions for the ultimate moment capacity of the reinforced beams were more conservative than those for the unreinforced beams of the same depth. As well as the reasons stated for the unreinforced beams, it is believed that if plasticization of the timber in the compression zone was taken into account in the analysis, agreement could be improved. Theoretical analyses were also undertaken using a design modulus of elasticity of 8000 N/mm 2 for the C16 grade timber in accordance with EN 338 and 45,000 N/mm 2 for the FRP plate in accordance with the manufacturer s literature. A design tensile strength of 12.6 N/mm 2, which included a lamination factor of 1.26, was used in the analyses. The design stiffness of the unreinforced beams (Phase A) is Nmm 2, FRP plate reinforced beams which do not include a sacrificial lamination (Phase B) is Nmm 2, unreinforced beams (Phase C) is Nmm 2 and FRP plate reinforced beams which include a sacrificial lamination (Phase D) is Nmm 2. The design ultimate moment capacities of the unreinforced beams (Phase A) is 7.28 knm, FRP plate reinforced which do not include a sacrificial lamination (Phase B) is 9.54 knm, unreinforced beams (Phase C) is 9.32 knm and FRP plate reinforced beams which include a sacrificial lamination (Phase D) is knm. It should be noted that the design strength and stiffness values used have not been modified to account for exposure class and duration of load effects. Further research is required to determine the appropriate modification factors. 7. Conclusions The flexural performance of unreinforced low-grade glulam and glulam which was reinforced with adhesively bonded recyclable FRP plates incorporating glass fibre reinforcement was compared. It was seen that by using practical reinforcement percentages and strategically locating the reinforcing material in the more highly stressed tensile region of the beam, pseudo-yielding of the compression fibres in the top timber lamination can occur. This failure behaviour introduces significant ductility and allows for a more efficient use of the timber in the compression laminations before tensile rupture occurs either at a knot or in clear wood at the bottom of the structural member. In order to enhance the