Basalt FRP rods for reinforcement and repair of timber

Composites: Part B 70 (2015) 9–19

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Composites: Part B journal homepage: www.elsevier.com/locate/compositesb

Basalt FRP rods for reinforcement and repair of timber Gary M. Raftery a,⇑, Fiona Kelly b a b

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

a r t i c l e

i n f o

Article history: Received 19 March 2014 Received in revised form 15 October 2014 Accepted 18 October 2014 Available online 30 October 2014 Keywords: A. Wood A. Polymer–matrix composites (PMCs) B. Strength D. Mechanical testing

a b s t r a c t Limited research has been undertaken into the use of basalt fibre reinforced polymer (FRP) materials for the strengthening and repair of structural timber elements. This paper describes an experimental test programme in which the flexural performance of low-grade glued laminated timber was reinforced using bonded-in basalt FRP rods. Tension test results show that basalt FRP rods compare extremely well to the mechanical characteristics of glass FRP rods. Strengthened and repaired beams exhibited considerable ductility in contrast to brittle tension behaviour of the unreinforced sections. With the use of a modest reinforcement percentage of 1.4% strategically located in circular routed out grooves at the soffit of the beam, mean stiffness enhancements of 8.4% and 10.3% for the global and local measurements were achieved respectively and a mean improvement in the ultimate moment capacity of 23% was achieved in comparison to the unreinforced glulam beams. The distance of the reinforcement to the neutral axis was shown to be highly influential on the mechanical enhancements. The use of basalt FRP rods is seen to be highly effective as a repair technology for damaged timber elements. Strain profile readings from the beams which included the reinforcement demonstrated improved utilisation of the compression characteristics of the timber. In all testing, a good quality bond is reported between the basalt FRP and wood. There exists significant potential for the development of environmentally friendly engineered structural elements by combining timber based products with other natural materials such as basalt fibre reinforced polymers. Ó 2014 Elsevier Ltd. All rights reserved.

1. General Introduction The use of basalt fibre reinforced polymers (FRP) in the construction industry is relatively new with limited research conducted with timber. The environmental benefits of the use of sustainable materials such as timber in the construction industry is documented in the literature [1]. There is, at present, significant interest in improving the mechanical performance of timber with the use of secondary materials. FRP rods have excellent mechanical properties, are lightweight and have good chemical and corrosion resistance [2]. The reinforcement of structural timber elements with a product manufactured from another natural material would be strongly favoured from an environmental aspect. Basalt is one of the most commonly occurring rock types and basalt fibres possess significantly lower global warming potential than steel and other synthetic fibres. This paper describes an experimental investigation that was undertaken to investigate the performance of

⇑ Corresponding author at: Department of Civil and Environmental Engineering, The University of Auckland, New Zealand. Tel.: +64 9 923 8536. E-mail address: [email protected] (G.M. Raftery). http://dx.doi.org/10.1016/j.compositesb.2014.10.036 1359-8368/Ó 2014 Elsevier Ltd. All rights reserved.

low-grade glued laminated timber strengthened and repaired using internally bonded basalt FRP rods.

1.1. Basalt FRPs in structural engineering The literature on the use of basalt fibres in structural engineering applications is limited. The use of basalt fibre as a strengthening material for concrete was experimentally examined in relation to durability, mechanical properties and flexural strengthening [3]. Basalt fibres have been used as reinforcement in concrete bridge deck slabs because of their improved corrosion resistance [4] and in geopolymeric concrete [5]. The confinement of concrete using basalt fibres bonded with a cement-based mortar was seen to overcome some limitations associated with epoxy based FRP laminates [6]. Hybrid glass-basalt FRP laminates performed equally as well as GFRP laminates when tested for column confinement [7]. Basalt fibre reinforced composites compared well to glass fibre reinforced composites when tested for corrosion resistance in seawater [8]. Debonding was delayed when near surface mounted basalt bars were contrasted with externally bonded reinforcement when strengthening concrete [9]. Basalt fibres may be considered as a

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possible alternative to glass in marine applications [10]. Basalt FRP strengthened beams demonstrated better performance in comparison to glass FRP when subjected to elevated temperatures [11]. Basalt textile-reinforced mortar layers provided substantial gains in the shear capacity of reinforced concrete beams [12]. Steel wire basalt fibre reinforced polymers were used to prolong the fatigue life of steel beams [13]. The number of studies undertaken using basalt fibre reinforcement in timber engineering is limited. Basalt fibre rods were used as internally bonded reinforcement for the strengthening of lowgrade glued laminated timber [14]. A pilot study was undertaken to assess the feasibility of strengthening timber with pre-stressed basalt fabric [15]. It was concluded that much further research was required. Basalt fabric was used with good results to reinforce solid wooden beams [16]. In more recent studies, the pull-out behaviour of bonded-in basalt fibre rods [17], and the reinforcement of short span beams with U shaped profiles of basalt fibre [18] were examined. A further study examined wood reinforcement with basalt, hemp and flax fibres in an epoxy resin and it was concluded that limited production costs, energy consumption and disposal at the end of their life were beneficial factors for such reinforcement [19]. 1.2. Strengthening and repair of timber elements with internal reinforcement Early experimental efforts involving the strengthening of timber elements using FRP composites are documented in the literature [20]. The reinforcement of low-grade glued laminated timber using recyclable FRP plates has been investigated both experimentally and with a nonlinear numerical model [21,22]. Limited published literature exists on the use of internal FRP reinforcements for the strengthening of low-grade wood. GFRP bars have been used to reinforce creosote treated timber [23] and have been used as shear reinforcement in timber elements [24]. Placement of the reinforcement at the neutral axis was not advised [25]. High strength steel fibres have been used to strengthen timber [26]. The geometry of the routed out grooves was shown to be influential when strengthening low-grade laminated spruce with glass FRP rods [27]. The use of several smaller diameter rods in the same groove proved to be of no advantage. A cost evaluation demonstrated that FRP reinforcement of timber was more appropriate to deeper beams [28] and promise existed with the use of low-quality timber [29]. Guidance regarding procedures relating to repair in timber structures is available in the literature [30,31]. Premature delamination was reported as a concern when rehabilitating timber sleepers with CFRP materials [32]. CFRP bars inserted into wood beams which were concealed from visibility were shown to be effective for repair applications [33]. FRP plates were used to repair horizontal splits with good success [34]. Steel reinforcement rods were seen to be the most effective in restoring stiffness when repairing small scale timber beams [35]. One hundred year old timber beams were repaired and tested with the wood FRP bond quality noted as being of high importance [36].

behaviour, failure mode, enhancements in both stiffness and ultimate moment capacity and strain profile distribution.

2. Materials 2.1. Timber Irish grown Sitka spruce was the timber species studied in this experimental test programme. This species is very fast growing because of the moist environment in Ireland. In order to assess the performance of the reinforcement as accurately as possible, all the timber used in the study was sourced in one delivery from the same sawmill. The timber was plain sawn cut, mechanically graded to C16 in accordance with EN 338 [37] and was kiln dried to approximately 18% moisture content. The boards were 4200 mm in length with a nominal section size of 96 mm  44 mm. The boards were conditioned to an environment of 65 ± 5% relative humidity and 20 ± 2 °C temperature once delivered to the laboratory. A mean equilibrium moisture content of 11.7% with standard deviation of 0.5% was obtained after the conditioning period and a mean board density of 392 kg/m3 with a standard deviation of 44 kg/m3 was recorded. 2.2. Basalt fibre rods The basalt FRP rods used as the reinforcement in this test programme were12 mm diameter Rockbar which were supplied by MagmaTech Ltd. The material has excellent chemical and corrosion resistance and has a coarse sanded surface finish which facilitates with mechanical interlocking at the adhesively bonded interface to the wood. Little waste material is produced during the manufacture of these rods and they are associated with considerably lower global warming impact potential in comparison to steel. Basalt FRP rods are also considerably lighter than steel. The mechanical properties of the rods in comparison to Glass FRP rods and Irish grown Sitka spruce are given in Table 1. Both modulus of elasticity given by the manufacturer and results as tested are presented. The basalt FRP rods compare well with the performance of the Glass FRP rods. 2.3. Adhesive The results obtained from previous research programmes which examined the ambient bond quality and durability quality of wood-laminating adhesives and epoxy adhesives for the bonding of FRP material to fast growing spruce were used to select appropriate adhesives for the current experimental test programme [41,42]. A phenol resorcinol formaldehyde (PRF) adhesive was selected for the bonding of the timber laminations and the wellrecognised civil engineering epoxy adhesive, Sikadur 31 was the

Table 1 Mechanical properties of basalt FRP, glass FRP and timber.

1.3. Objectives of the present study The objective of this research is to examine the use of bonded-in basalt FRP rods for the strengthening and repair of timber elements. The experimental test programme involved the fabrication and testing in flexure of unreinforced, strengthened, artificially fractured and repaired glulam beams. The mechanical performance of the unreinforced, reinforced, artificially fractured and repaired beams is compared with regard to the load–deflection

a

Material

Ultimate tensile strength (N/mm2)

Modulus of elasticity in tension (N/mm2)

Basalt FRP rods

1000+a

Glass FRP rods Wood

620c 23.7e

45,000+a 49,000–51,000b 47,000–52,000d 8111e

Basalt FRP rod properties as reported by manufacturer [38]. Basalt FRP rod properties as tested from ten number specimens. Glass FRP rod properties as reported by manufacturer [27]. d Glass FRP rod properties as tested [27]. e In-grade testing with mean moisture content of 12% and mean density of 403 kg/m3 [39,40]. b c

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adhesive which was deemed most suitable for the bond between the basalt reinforcing rods and wood. 3. Experimental testing 3.1. Test programme The test programme involved the fabrication and testing of unreinforced, strengthened, artificially fractured and repaired glulam beams. Stiffness testing and testing to failure was undertaken for each beam. The unreinforced beams were tested in order to act as a control for the strengthened, artificially fractured and repaired beams. The artificially fractured beams involved a saw kerf through the depth of the bottom lamination at midspan which simulated a damaged member. Two reinforcement and repair arrangements using tension only reinforcement were studied. The reinforcement was placed in routed out grooves in the soffit of the beams or in routed out grooves on the sides of the beams. The reinforcement percentage studied was 1.4%. The beam section configurations are shown in Fig. 1 and the test programme is shown in Table 2. In order to ensure confidence in the results, five repetitions were tested for each configuration. All dimensions shown in the figures are in millimetres. 3.2. Manufacture of beams The lamination stock was initially assessed using a three point bending mechanical stress grader. The stock was subsequently assessed visually and boards which had excessive warping were removed. Strength reducing defects were assessed for all boards. The beams were manufactured from five number 38 mm thick laminations bonded together to give an end length of 3610 mm and beam depth of 190 mm. A beam span of 3420 mm was used. This dimension gave a span-depth ratio of 18:1 as recommended for the four point bending test specified in EN 408 [43]. The manufacturing procedure for the beams involved strategic balanced lay-ups using the best quality laminations in the most extremely stressed tension zone at the bottom of each beam. Beams with a

Table 2 Experimental test programme. Phase

Description

Repetitions

Beam numbers

A B C D E

Unreinforced Reinforced at soffit Reinforced at sides Artificially fractured Repaired artificially fractured at soffit Repaired artificially fractured at sides

5 5 5 5 5

1, 7, 2, 8, 3, 9, 4, 10, 5, 11,

5

6, 12, 18, 24, 30

F

13, 14, 15, 16, 17,

19, 20, 21, 22, 23,

25 26 27 28 29

lower beam number were theoretically stiffer and stronger than beams of higher number. The timber laminations were bonded together using a PRF adhesive. The orientations of the annular rings in the laminations during manufacture complied with the requirements as specified in EN 386 [44]. An adhesive spread rate of 400 g/ m2 was used. A pressure of 0.7 N/mm2 was applied to the closed assembly for 24 h in an environment of 65 ± 5% relative humidity and 20 ± 2 °C temperature. Circular shaped grooves were routed into the beams using a computer numerically controlled machine to facilitate placement of the basalt FRP reinforcement. The size of the grooves allowed for a 2 mm adhesive bondline. Surface preparation of the basalt FRP rods involved removing any dust or residues by means of wiping with a dry cloth initially and subsequently wiping clean using methylated spirits. No longer than three hours was allowed to elapse prior to application of the epoxy adhesive. The beams were stored in an environmental conditioning chamber with conditions of 65 ± 5% relative humidity and 20 ± 2 °C temperature for a period of at least 30 days prior to testing to ensure the adhesive was fully cured.

3.3. Testing of beams All beam testing was undertaken in accordance with the four point bending arrangement as specified in EN 408 [43]. The test arrangement is illustrated in Fig. 2. The beams were initially tested in their unreinforced state and later again in their reinforced, frac-

Fig. 1. Beam section configurations (a) Phase A/Phase C (b) Phase B/Phase D (c) Phase C/Phase E.

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Fig. 2. Stiffness test arrangement.

tured or repaired state so that a direct comparison between the unreinforced and reinforced or repaired strategies could be obtained. Both the local and global stiffness was measured. An inverted linear variable differential transformer (LVDT), located at midspan is used to measure the global stiffness of each beam. A second LVDT which is positioned in a hanger suspended from the neutral axis over a distance of five times the depth of the beam is used to measure the local stiffness (Fig. 2). A preload of 300N was applied to all beams prior to a displacement rate of 0.57 mm/s being used for the stiffness testing. Steel plates of 10 mm thickness, 95 mm length and 90 mm width were used at the loading heads and supports to minimise local indentations. It was ensured that the load did not exceed the elastic limit or 40% of the ultimate load throughout the stiffness testing of any of the beams. Lateral restraint to all the beams was positioned approximately 300 mm outside the loading heads. Two polytetrafluoroethylene (PTFE) strips were allowed to slide over each other at the lateral supports so that the effects of friction were minimised. The local stiffness LVDT and associated hanger was removed prior to loading to failure. The strain distribution behaviour was recorded on the unreinforced Beam 13 (Phase A), and the reinforced beams, Beam 14 (Phase B) and Beam 15 (Phase C). These beams were selected as they were in the midrange of the beams manufactured and were most representative of the beams in each of the beam phases instrumented. A true representation would not have been achieved if the strongest or weakest beams were instrumented. The strain instrumentation involved positioning 60 mm long gauges (PL-60-11) on both sides at midspan throughout the depth of each beam. By placing gauges on either side of the beams, any twist that might occur was accounted for. The 60 mm gauge length reduced the influence associated with deviations that may occur from any localised irregularities along the grain. The arrangement used for the strain gauges as shown for the Phase A beam is shown in Fig. 3.

Fig. 3. Strain gauge arrangement for Phase A beam.

4. Experimental results and discussion 4.1. Load–deflection behaviour 4.1.1. Load–deflection behaviour, Phase A – unreinforced beams The Phase A beams comprised the unreinforced control beams of which their load–deflection behaviour to failure are shown in Fig. 4. Two of the beams exhibited limited nonlinear behaviour (Beam 1 and Beam 19). It should be noted that from the fabrication process which was undertaken, the Phase A beams theoretically had the best quality timber laminations and Beam 1 was theoretically the strongest beam of all the beams manufactured in the entire test programme when in its unreinforced state. Furthermore, the highest quality lamination was used at the bottom of Beam 1. This therefore can explain why this unreinforced beam exhibits the greater nonlinear behaviour. In Beam 19, the knot at which failure occurred was outside the zone of maximum bending moment. Therefore, weak areas within the zone of maximum bending moment were not weak enough to initiate failure before the beam experienced limited nonlinear behaviour. The remaining three beams (Beam 7, Beam 13 and Beam 25) all exhibited linear elastic behaviour and failed catastrophically. There was no visible compression wrinkling in the top laminations of any of the Phase A beams. In general, for timber where defects are present, the compressive strength will be greater than the average tensile strength [37]. Because of the low quality of timber being used, all the beams failed at defects or irregularities in the tension zone as the yield stress in tension was exceeded before the yield stress in compression was exceeded. 4.1.2. Load–deflection behaviour, Phase B – tension reinforcement using 12 mm diameter rods at soffit In contrast with the load–deflection behaviour demonstrated by the Phase A beams, significant nonlinear behaviour is experienced when the glued laminated beams are reinforced with the basalt FRP rods. Furthermore, despite the Phase B beams being manufactured with laminations of theoretically lower strength, there is a significant improvement in their load carrying capacities when compared to the Phase A beams, as illustrated in Fig. 5. Beam 8 failed in compression with deep compression wrinkles visible in the top laminations. All of the other Phase B beams fractured at defects or irregularities in the most highly stressed fibres of the bottom laminations after exhibiting nonlinear load–deflection behaviour. With the exception of Beam 8, the beams showed reserve load carrying capacity after this initial tensile fracture and deep compression wrinkles were seen to propagate in the top laminations as can be seen in Fig. 6 for Beam 2. At no stage did the reinforcement prematurely debond and cracking along the reinforcement adhesive interface only initiated after the tensile fractures occurred as can be seen in Fig. 7 for Beam 2. From visual inspection of the samples, the degree of ductility which was experienced prior to the initial fracture was directly dependent on the presence of defects and irregularities in the bottom laminations of the beams as some defects were removed when the grooves

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

Fig. 7. Fracture at defect in Beam 2 (Phase B).

Fig. 5. Load vs. deflection behaviour for reinforced beams at soffit (Phase B).

Fig. 8. Load vs. deflection behaviour for reinforced beams at sides (Phase C).

Fig. 6. Compression buckling in Beam 2 (Phase B).

were routed out and their presence replaced by the basalt FRP rods and epoxy adhesive. During testing it was also deemed that the reinforcement acted as a bridge across knots and strength reducing defects in the critical tension zone of the timber. 4.1.3. Load–deflection behaviour, Phase C – tension reinforcement using 12 mm diameter rods at sides An alternative configuration of strengthening was investigated for the Phase C beams in which the basalt reinforcement was bonded into routed out grooves in the sides of the beams. This configuration allowed for the reinforcement to be less visible from beneath the beams and hence preserve the aesthetic characteristics of the timber. The load–deflection behaviour of these beams is shown in Fig. 8. The beams are associated with a lower load

carrying capacity in comparison to the Phase B beams as the position of the reinforcement in relation to the neutral axis is not optimised and the reinforcement is above the most highly stressed tension fibres. All the beams exhibited similar load–deflection behaviour whereby the beams initially exhibited nonlinear behaviour before the wood below the reinforcement fractured at a defect. The basalt reinforcement remained intact and the load– deflection behaviour continued nonlinearly as micro-cracking propagated in the bottom timber laminations and compression wrinkles became clearly visible in the top laminations. The rupturing of the wood below the reinforcement is illustrated in Fig. 9 for Beam 21. Similar to the Phase B beams, no issues with premature debonding of the reinforcement were noted. 4.1.4. Load–deflection behaviour, Phase D – artificially fractured beams The load–deflection behaviour of the artificially fractured Phase D beams is shown in Fig. 10. These beams had a saw kerf at midspan through the bottom tension lamination to simulate a damaged element that required repair. The beams were tested to failure to examine the effect of the artificial fracture on the mechanical properties of the section. No compressive wrinkling at knots in the top laminations was recorded for any of the beams. All beams fractured prematurely at the artificial saw kerf without prior warning. None of the beams experienced an initial fracture at a defect away from the artificial fracture. After the initial fracture, the load–deflection behaviour of the beams became nonlinear as the crack propagated in the clear wood surrounding the location

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Fig. 9. Fracture at defect below reinforcement in Beam 21 (Phase C).

Fig. 12. Fracture at knot in Beam 29 (Phase E).

trates that the basalt FRP rods act as a bridge across the artificial fracture and are effective as a sustainable repair technology for timber structures. The load carrying capacity of the repaired beams is considerable in comparison to the artificial fractured Phase D beams. Cracking at the adhesive interface became evident only after the fractures in the timber were considerable. No issues with premature failure of the bond were recorded.

Fig. 10. Load vs. deflection behaviour for artificially fractured beams (Phase D).

of the artificial fracture. Beam 4 experienced limited nonlinear behaviour post the initial crack at the artificial fracture as a result of a large defect adjacent to the location. 4.1.5. Load–deflection behaviour, Phase E – artificially fractured beams repaired using 12 mm diameter rods at soffit As can be seen from the load–deflection curves shown in Fig. 11, non-linear behaviour is experienced in the Phase E beams. These beams had an artificial fracture in the bottom lamination and were repaired at the soffit using the basalt FRP rods. The load–deflection behaviour is similar to the Phase B beams although cracks initiate at lower loads near the artificial fracture. Compression wrinkling at knots in the top lamination was evident in all of the beams in Phase E. In one of the beams, the first fracture location initiated at defects in the bottom tension lamination away from the location of the artificial fracture as can be seen in Fig. 12. This behaviour illus-

Fig. 11. Load vs. deflection behaviour for artificially fractured repaired beams at soffit (Phase E).

4.1.6. Load–deflection behaviour, Phase F – artificially fractured beams repaired using 12 mm diameter rods at sides The load–deflection behaviour of the repaired Phase F beams is illustrated in Fig. 13. All the beams initially fractured in the wood of the bottom tension lamination below the reinforcement at the location of the artificial fracture. These fractures occurred at much higher load levels than the initial fractures which occurred in the Phase D beams and this indicates the effectiveness of the basalt FRP rods for repair purposes. The reinforcement remained in position despite the rupturing of the wood below. All the beams exhibited compressive wrinkling in the top compression lamination. In the Phase F beams, the reinforcement was positioned in grooves routed into the sides of the beams so that it would be less visible and more aesthetic than if the reinforcement was positioned in grooves at the soffit as in the Phase E beams. However, because the reinforcement is closer to the neutral axis in the Phase F beams and is above the more highly stressed wood fibres, fracture initiates at lower loads and the load carrying capacity of the repaired Phase F beams is lower than that of the Phase E beams. 4.2. Stiffness testing 4.2.1. Stiffness testing, Phase A – unreinforced beams The reference or control beams in the test programme were the unreinforced Phase A beams. These beams were tested for stiffness prior to being tested to failure. The beams were associated with a mean global stiffness of 5.45  1011 Nmm2 and standard deviation of 0.3  1011 Nmm2 and a mean local stiffness of 6.31  1011 Nmm2 and standard deviation of 0.7  1011 Nmm2 as shown in Fig. 14. The global stiffness measurements are lower as shear deformation is associated with the reading as well as there is a possibility of indentation at the supports because of the low density of the wood. 4.2.2. Stiffness testing, Phase B – tension reinforcement using 12 mm diameter rods at soffit The stiffness of the Phase B beams was measured both in the unreinforced and reinforced states and the results are illustrated in Fig. 15. In their unreinforced state, the beams had a mean global stiffness of 5.16  1011 Nmm2 with a standard deviation of 0.3  1011 Nmm2 and a mean local stiffness of 5.71  1011 Nmm2 with

G.M. Raftery, F. Kelly / Composites: Part B 70 (2015) 9–19

Fig. 13. Load vs. deflection behaviour for artificially fractured repaired beams at sides (Phase F).

a standard deviation of 0.62  1011 Nmm2. After the reinforcement was fitted, the beams had a mean global stiffness of 5.59  1011 Nmm2 with a standard deviation of 0.31  1011 Nmm2 and a mean local stiffness of 6.28  1011 Nmm2 with a standard deviation of 0.54  1011 Nmm2. These values represent increases of 8.4% and 10.3% in the mean global stiffness and mean local stiffness. The local stiffness measurements are higher because shear deflection is not included in these readings. The possibility of indentation at the supports was reduced when the beams were tested with the reinforcement at the bottom. 4.2.3. Stiffness testing, Phase C – tension reinforcement using 12 mm diameter rods at sides Testing was undertaken both in the unreinforced and reinforced states for the Phase C beams. The results are shown in Fig. 16. When unreinforced, the beams had a mean global stiffness of 5.14  1011 Nmm2 with a standard deviation of 0.37  1011 Nmm2 and a mean local stiffness of 5.8  1011 Nmm2 with a standard deviation of 0.39  1011 Nmm2. After the reinforcement procedure and the beams were retested, the Phase C beams had a mean global stiffness of 5.43  1011 Nmm2 with a standard deviation of 0.4  1011 Nmm2 and a mean local stiffness of 6.15  1011 Nmm2 with a standard deviation of 0.5  1011 Nmm2. This corresponds to a mean global stiffness increase of 5.8% and a mean local stiffness increase of 5.7%. Comparing this performance to that of the Phase B beams, the importance of optimising the position of the reinforcement in the section is highlighted. 4.2.4. Stiffness testing, Phase D – artificially fractured beams The results from the testing with the Phase D beams are illustrated in Fig. 17. The beams had a mean global stiffness of 5.13  1011 Nmm2 with a standard deviation of 0.22  1011 Nmm2 and a mean local stiffness of 5.87  1011 Nmm2 with a standard

Fig. 14. Stiffness testing. Phase A: Beams 1, 7, 13, 19 and 25. Key: A = Global stiffness unreinforced beams; B = Local stiffness for unreinforced beams.

15

Fig. 15. Stiffness testing. Phase B: Beams 2, 8, 14, 20 and 26. Key: A = Global stiffness for beams when unreinforced; B = Local stiffness for beams when unreinforced; C = Global stiffness for reinforced beam at soffit; D = Local stiffness for reinforced beam at soffit.

deviation of 0.35  1011 Nmm2 when unreinforced. After the artificial fracture was placed in the bottom tension lamination the beams had a mean global stiffness of 4.29  1011 Nmm2 with a standard deviation of 0.38  1011 Nmm2 and a mean local stiffness of 3.1  1011 Nmm2 with a standard deviation of 0.42  1011 Nmm2. This represents a 16.3% reduction in the global stiffness and a 47% reduction in the local stiffness. The effect of the artificial fracture is considerably greater for the local stiffness as this reading is taken in the zone of maximum bending moment where the fracture is located.

4.2.5. Stiffness testing, Phase E – artificially fractured beams repaired using 12 mm diameter rods at soffit The Phase E beams were stiffness tested initially when unreinforced, then after artificial fractures were placed in the bottom lamination of each beam and, lastly, after each beam was repaired with basalt FRP rods bonded into grooves at the soffit. The beams when unreinforced had a mean global stiffness of 5.37  1011 Nmm2 with a standard deviation of 0.1  1011 Nmm2 and a mean local stiffness of 6.42  1011 Nmm2 with a standard deviation of 0.54  1011 Nmm2. When the fractures were placed in the beams, the global and local stiffness values were reduced by 19% and 50.4%, respectively. After the repair procedure using the basalt FRP rods was completed, the beams had a global stiffness of 5.65  1011 Nmm2 with a standard deviation of 0.3  1011 Nmm2 and a mean local stiffness of 5.98  1011 Nmm2 with a standard deviation of 0.6  1011 Nmm2. In comparison to the original stiffness properties of the beams, this corresponded to a mean

Fig. 16. Stiffness testing. Phase C: Beams 3, 9, 15, 21 and 27. Key: A = Global stiffness for beams when unreinforced; B = Local stiffness for beams when unreinforced; E = Global stiffness for reinforced beam at sides; F = Local stiffness for reinforced beam at sides.

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Fig. 17. Stiffness testing. Phase D: Beams 4, 10, 16, 22 and 28. Key: A = Global stiffness for beams when unreinforced; B = Local stiffness for beams when unreinforced; G = Global stiffness for artificially fractured beam; H = Local stiffness for artificially fractured beam.

5.1% increase in global stiffness and a mean 6.4% reduction in local stiffness. The results are illustrated in Fig. 18. 4.2.6. Stiffness testing, Phase F – artificially fractured beams repaired using 12 mm diameter rods at sides The Phase F beams were stiffness tested after fabrication in their unreinforced state, tested again after the artificial fracture was placed in the bottom lamination and tested a third time after the beams were repaired at the sides using bonded-in basalt FRP rods. When unreinforced, the beams had a mean stiffness of 5.35  1011 Nmm2 and mean global stiffness of 5.76  1011 Nmm2 with standard deviations of 0.2  1011 Nmm2 and 0.54  1011 Nmm2, respectively. After the artificial fracture was placed in the beams the mean global and local stiffness was measured as 4.23  1011 Nmm2 and 3.16  1011 Nmm2, respectively, which represented a mean reduction of 20.9% and 44.7% in comparison to the original undamaged sections. When the beams were repaired, the mean global stiffness was recorded as 5.22  1011 Nmm2 and the mean local stiffness as 5.35  1011 Nmm2, as shown in Fig. 19. These values correspond to a 2.3% and 7.5% reduction in relation to the original stiffness performance of the unreinforced Phase F beams and demonstrate the effectiveness of the basalt FRP rods for repairs relating to deflection concerns when aesthetics are a concern. 4.3. Ultimate moment capacity The results from the failure tests in relation to ultimate bending moment capacity Mult for each of the beam phases are shown in

Fig. 19. Stiffness testing. Phase F: Beams 6, 12, 18, 24 and 30. Key: A = Global stiffness for beams when unreinforced; B = Local stiffness for beams when unreinforced; G = Global stiffness for artificially fractured beam; H = Local stiffness for artificially fractured beam; K = Global stiffness for repaired artificially fractured beam at sides; L = Local stiffness for repaired artificially fractured beam at sides.

Table 3. The Phase B beams reinforced at the soffit exhibited the most significant performance enhancement with an increase of 23.1% when compared to the unreinforced control Phase A beams. The average increase for when the beams were reinforced at the sides (Phase C) was 12.1%. The primary difference between the Phase B and Phase C beams was the depth of the reinforcement in the section. The reinforcement percentage remained the same at 1.4%. The artificial fracture through the bottom lamination at midspan in the Phase D beams had a significant effect on the ultimate moment capacity of the beams with a 60.8% reduction in comparison to the performance of the unreinforced control Phase A beams. When the artificially fractured beams were repaired at the soffit with the basalt FRP rods as for the Phase E beams, their performance easily surpassed the performance of beams in their unreinforced state. These repaired beams achieved a mean ultimate moment capacity of 27.6 kNm which represented a mean increase of 18.4% when compared to the Phase A beams demonstrating the effectiveness of the repair strategy. When the beams were repaired with the reinforcement positioned in the routed out grooves at the sides (Phase F), the beams achieved a mean ultimate moment capacity of 21.2 kNm which was a significant improvement in relation to the mean performance of 9.2 kNm for the ultimate moment capacity of the Phase D fractured beams. However, this repair strategy did not succeed in reaching the mean ultimate moment capacity of the Phase A beams as the mean performance of the Phase F was approximately 9% lower. Nevertheless, it should be considered that the beams tested are of a prototype size and in deeper sections the reinforcement would be further from the neutral axis and hence the effect of the basalt FRP rods would be enhanced. Furthermore, the Phase F beams were manufactured with theoretically the lowest quality timber for all the beam phases tested and therefore, these beams would theoretically fail at the lowest ultimate moment capacities if tested in their unreinforced state. There was no significant difference between the variability of the results determined for the various beam phases.

4.4. Strain profile distribution

Fig. 18. Stiffness testing. Phase E: Beams 5, 11, 17, 23 and 29. Key: A = Global stiffness for beams when unreinforced; B = Local stiffness for beams when unreinforced; G = Global stiffness for artificially fractured beam; H = Local stiffness for artificially fractured beam; I = Global stiffness for repaired artificially fractured beam at soffit; J = Local stiffness for repaired artificially fractured beam at soffit.

The Moment vs. Strain behaviour for Beam 13 is shown in Fig. 20. Beam 13 was instrumented as theoretically, it would give the best indication of the performance of the unreinforced Phase A beams as this beam was manufactured from middle of the range quality laminations in the timber stock. Deviations between readings from gauges on opposite sides of the beam resulted because of the non-homogeneous nature of the timber as well as any twist in

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G.M. Raftery, F. Kelly / Composites: Part B 70 (2015) 9–19 Table 3 Ultimate bending moments. Phase A

Mult (kNm)

Phase B

Mult (kNm)

Phase C

Mult (kNm)

Phase D

Mult (kNm)

Phase E

Mult (kNm)

Phase F

Mult (kNm)

Beam 1 Beam 7 Beam 13 Beam 19 Beam 25 Mean St. dev.

25.2 23.3 19.9 23.9 24.2 23.3 2.0

Beam Beam Beam Beam Beam

30.2 29.7 27.2 30.6 25.8 28.7 2.1

Beam Beam Beam Beam Beam

24.8 28.2 27.7 24.0 26.5 26.2 2.1

Beam Beam Beam Beam Beam

5.8 9.8 10.7 12.1 7.9 9.2 2.5

Beam Beam Beam Beam Beam

26.7 28.3 26.9 31.0 25.2 27.6 2.2

Beam Beam Beam Beam Beam

22.1 20.4 20.3 22.0 21.1 21.2 0.8

2 8 14 20 26

3 9 15 21 27

4 10 16 22 28

5 11 17 23 29

6 12 18 24 30

Fig. 20. Moment vs. Strain for Beam 13 (Phase A).

the beam. Limited nonlinear behaviour is recorded by the gauges in the compression zone. There was no evidence of compression wrinkling and the maximum strain recorded by SG1, the strain gauge at the top of the beam, was 3706  16. Strain behaviour in the tension zone is predominantly linear up to an applied moment of 17 kNm, approximately. Visible cracking initiated at a defect in the bottom tension lamination at this load level. This cracking was followed by a series of further micro-cracks as the applied load was increased. Non-linear behaviour is recorded by the gauges in the tension zone as a result of these fractures. The strain profile of the beam, as shown in Fig. 21, illustrates that when a moment of 10 kNm is applied, the neutral axis is approximately 2.2% lower than the centroid of the section. The neutral axis is lower than the centroid as the manufacturing procedure for all the beams in the test programme involved positioning the best quality laminations in the most highly stressed tension zone at the bottom of the beam. The Moment vs. Strain plots for the instrumented Phase B beam, Beam 14, is shown in Fig. 22. This beam phase included the basalt FRP rod reinforcement positioned in routed out grooves at the bottom of the beams. The strain gauge, SG1, which was located at midspan on the top lamination did not record readings from the mostly highly strained localised zone as this occurred at a defect adjacent to the location of the gauge. However, significant

Fig. 21. Strain profile for Beam 13 (Phase A).

Fig. 22. Moment vs. Strain for Beam 14 (Phase B).

strain behaviour was recorded by one of the strain gauges on the side of the top compression lamination, SG3, as a compression wrinkle initiated at a defect in the zone of the gauge. Nonlinear strain behaviour is recorded by the gauges as low as the centre lamination in the beam (SG6 and SG7) which demonstrates that increased ductility can be associated with beams which include the Basalt FRP reinforcement. Strain behaviour to failure in the tension zone, is largely seen to be elastic with the exception of some micro-cracking at high moment levels. Such micro-cracks initiated at defects in the 16 mm thick wooden ridge to the outside of the grooves in which the reinforcement was positioned. The strain profile for Beam 14 is shown in Fig. 23. At an applied moment of 10 kNm, the neutral axis is approximately 6.6% lower when compared to the depth of the centroid in the section. This is a significant difference when compared to the Phase A beam. As the applied moment approaches the ultimate moment capacity of 27.19 kNm, it can be seen that the neutral axis deepens in the section as a result of plasticization in the compression zone.

Fig. 23. Strain profile for Beam 14 (Phase B).

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G.M. Raftery, F. Kelly / Composites: Part B 70 (2015) 9–19

Fig. 25. Strain profile for Beam 15 (Phase C). Fig. 24. Moment vs. Strain for Beam 15 (Phase C).

The Moment vs. Strain plot for Beam 15 (Phase C) which included basalt FRP rod reinforcement at the sides is shown in Fig. 24. Despite Beam 15 being manufactured from laminations which theoretically represented the mean performance of the Phase C beams, Beam 15 exhibited the highest load carrying capacity of the Phase C beams prior to fracture and had the second highest ultimate moment capacity of the beams in the phase. The strain gauges SG1, SG2 and SG3 failed to record the maximum strains in the top laminate towards ultimate moment capacity as significant compression wrinkles initiated at a knot away from the midspan. At a moment of 10 kNm, SG1 recorded a compressive strain of 1207  16 on Beam 15 in comparison to a strain of 1906  16 when the same moment level was applied on Beam 14 (Phase B). The higher strain reading associated with Beam 14 illustrates the effectiveness of positioning the reinforcement at the optimum distance from the neutral axis. In Beam 15, the reinforcement is positioned in grooves routed into the sides in order to conceal it from beneath the beam. Similar to the behaviour of Beam 14 (Phase B), nonlinear compressive strain behaviour is recorded by gauges as low as those on the centre lamination in the beam, SG6 and SG7. The strain behaviour recorded by gauges in the tension zone is linear up to an applied moment of 22 kNm, approximately. Bilinear behaviour is experienced at higher applied moments. Such behaviour occurred as the bottom tension lamination fractured at a knot and the fracture propagated along the grain towards the end of the beam. This behaviour reduced the second moment of area of the beam and the stiffness of the section. The position at midspan where the strain gauge was bonded to the bottom lamination remained unaffected by the fracture and the strain gauge continued to measure readings as the moment was increased. Differences between readings from gauges on opposite sides of the beam were again believed to result from variations in characteristics of the wood. The strain profile shown in Fig. 25 indicates that Beam 15 does not appear to utilise the compressive characteristics of the timber as effectively as Beam 14. At an applied moment of 10 kNm, the neutral axis in Beam 15 is at a lower depth than in Beam 14 as the reinforcement is much closer the neutral axis. Similar to the behaviour shown by Beam 14, as the applied moment approaches ultimate moment capacity and plastic compressive behaviour occurs in the top laminations, the neutral axis deepens. The position of the neutral axis at ultimate moment capacity compares well to the performance of Beam 14 (Phase B). This is because Beam 15 was one of the best performing Phase C beams. The ultimate moment capacity of Beam 15 was 27.7 kNm in comparison to 27.19 kNm for Beam 14. A further reason to justify why the depth of the neutral axis in Beam 15 compares well to the depth in Beam 14 would be that the laminations in Beam 15 are theoretically manufactured with less stiff laminations. The use of

the reinforcement rods in the tension zone would deepen the neutral axis more significantly if lower quality timber was being used. 5. Conclusions The results of a test programme to examine the use of basalt fibre reinforced polymer rods for the strengthening and repair of low-grade glued laminated timber in flexure was discussed. Several important observations were made:  Mechanical properties of basalt FRP rod compare favourably to the properties of glass FRP rod as illustrated in Table 1 with regard to strengthening and repair of timber.  For strengthening applications, using a modest 1.4% reinforcement, the basalt FRP rods can enhance stiffness by over 10% and ultimate moment capacity over 23%.  For repair applications, the basalt FRP rods can effectively bridge over damaged zones in the timber and comprehensively restore the mechanical strength and stiffness of the original undamaged section.  Unreinforced beams generally exhibited brittle tensile failures with fractures initiating at defects in the bottom laminations. Reinforced and artificially fractured repaired beams which used bonded-in basalt FRP rods in the tension zone generally exhibited considerable ductility with visible compression wrinkling in the top laminations.  The amount of ductility that was experienced by the reinforced sections was influenced by the distance of the reinforcing rods from the neutral axis.  Strain profile readings indicated that greater utilisation of the nonlinear compressive characteristics of the timber are achieved in the reinforced beams.  No issues were evident in relation to the integrity of the bond between the wood and basalt FRP rods.  With increasing acceptance of basalt FRP profiles in the construction industry, there exists considerable potential for the development of sustainable basalt-timber hybrid elements.

Acknowledgments This research was undertaken at the National University of Ireland, Galway. The primary author would like to express sincere thanks to the National University of Ireland, Galway for the financial support provided. The allocation of research funding from the Vocational Education Committee (Ireland) to support the postgraduate research student, Ms. Fiona Kelly, is acknowledged. The basalt FRP material provided by MagmaTech Ltd. is greatly appreciated.

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References [1] Khatib JM. Sustainability of construction materials. Cambridge (UK): Woodhead Publishing Limited; 2009. [2] Hollaway LC, Teng JG. Strengthening and rehabilitation of civil infrastructures using fibre-reinforced polymers (FRP) composites. Cambridge (UK): Woodhead Publishing Limited; 2008. [3] Sim J, Park C, Young Moon D. Characteristics of basalt fiber as a strengthening material for concrete structures. Compos Part B Eng 2005;36:504–15. [4] Tharmarajah G, Taylor SE, Robinson D, Cleland DJ. Composite reinforcement for bridge decks. In: Proceedings of the symposium of bridge and infrastructure and concrete research in Ireland, (BCRI 2010), Cork, Ireland, 2010, p. 125–132. [5] Li W, Xu J. Mechanical properties of basalt fiber reinforced geopolymeric concrete under impact loading. Mater Sci Eng A 2009;505:178–86. [6] Di Ludovico M, Prota A, Manfredi G. Structural upgrade using basalt fibers for concrete confinement. J Compos Constr 2010;14(5):541–52. [7] De Luca A, Nardone F, Matta F, Nanni A, Lignola GP, Prota A. Structural evaluation of full-scale FRP confined reinforced concrete columns. J Compos Constr 2011;15(1):112–23. [8] Wei B, Cao H, Song S. Degradation of basalt fibre and glass fibre/epoxy resin composites in seawater. Corros Sci 2011;53:426–31. [9] Bilotta A, Cerona F, Di Ludovico M, Nigro E, Pecce M, Manfredi G. Bond efficiency of EBR and NSM FRP systems for strengthening concrete members. J Compos Constr 2011;15(5):757–72. [10] Fiore V, Di Bella G, Valenza A. Glass-basalt/epoxy hybrid composites for marine applications. Mater Des 2011;32:2091–9. [11] Tan KH, Zhou Y. Performance of FRP strengthened beams subjected to elevated temperatures. J Compos Constr 2011;15(3):304–11. [12] Al-Salloum YA, Elsanadedy HM, Alsayed SH, Iqbal RA. Experimental and numerical study for the shear strengthening of reinforced concrete beams using textile reinforced mortar. J Compos Constr 2012;16(1):74–90. [13] Wu G, Wang HT, Wu ZS, Liu HY, Ren Y. Experimental study on the fatigue behaviour of steel beams strengthened with different fiber-reinforced composite plates. J Compos Constr 2012;16(2):127–37. [14] Kelly F. Basalt FRP rods for strengthening and repair of glued laminated timber [MEngSc thesis]. National University of Ireland, Galway, Ireland; 2012. [15] Zachary C, Kavan S. Feasibility of strengthening glulam beams with prestressed basalt fibre reinforced polymers [Master’s thesis]. Chalmers University of Technology, Göteborg, Sweden; 2012. [16] Speranzini E, Agnetti S. Structural performance of natural fibers reinforced timber beams. In: Proc. 6th international conf on compos in civil eng., June 2012. Rome, Italy: Paper 01-693, 2012. [17] Yeboah D, Taylor S, McPolin D, Gilfillan R. Pull-out behaviour of axially loaded basalt fibre reinforced polymer (BFRP) rods bonded perpendicular to the grain of glulam elements. Constr Build Mater 2013;38:962–9. [18] de la Rosa García P, Cobo Escamilla A, Nieves González Garcia M. Bending reinforcement of timber beams with composite carbon fiber and basalt fiber materials. Compos Part B Eng 2013;55:528–36. [19] Borri A, Corradi M, Speranzini E. Reinforcement of wood with natural fibres. Compos Part B Eng 2013;53:1–8. [20] Bulleit WM. Reinforcement of wood materials: a review. Wood Fiber Sci 1984;16(3):391–7. [21] Raftery GM, Harte AM. Low-grade glued laminated timber reinforced with FRP plate. Compos Part B Eng 2011;42(4):724–35.

19

[22] Raftery GM, Harte AM. Nonlinear numerical modelling of FRP plate reinforced glued laminated timber. Compos Part B Eng 2013;52:40–50. [23] Gentile C, Svecova D, Rizkalla SH. Timber beams strengthened with GFRP bars: development and applications. J Compos Constr 2002;6(1):11–20. [24] Svecova D, Eden RJ. Flexural and shear strengthening of timber beams using glass fibre reinforced polymer bars in experimental investigation. Can J Civ Eng 2004;31:41–55. [25] Ribeiro AS, de Jesus AMP, Lima AM, Lousada JLC. Study of strengthening solutions for glued laminated wood beams of maritime pine wood. Constr Build Mater 2009;23:2738–45. [26] Borri A, Corradi M. Strengthening of timber beams with high strength steel cords. Compos Part B Eng 2011;42(6):1480–91. [27] Raftery GM, Whelan C. Low-grade glued laminated timber beams reinforced using improved arrangements of bonded-in GFRP rods. Constr Build Mater 2014;52:209–20. [28] Stevens ND, Criner GK. Economic analysis of fiber-reinforced polymer wood beams. Maine agriculture and forest experimentation station bulletin, vol. 848. University of Maine; 2000. [29] Martin ZA, Stith JK, Tingley DA. Commercialisation of FRP reinforced glulam beam technology. In: Proceedings of the 6th world conference on timber engineering, Whistler; 2000. [30] Van Gemert D, Vanden Bosch M. Structural restoration of wooden beams by means of epoxy resin. Mater Struct 1987;20:165–70. [31] Wheeler AS, Hutchinson AR. Resin repairs to timber structures. Int J Adhes Adhes 1998;18(1):1–13. [32] Kropf FC, Meierhofer U. Strengthening, retrofitting and upgrading of timber structures with high-strength fibres. Struct Eng Int 2000;10(3):178–81. [33] Radford DW, Van Goethem D, Gutkowski RM, Peterson ML. Composite repair of timber structures. Constr Build Mater 2002;16(7):417–25. [34] Borri A, Corradi M, Grazini A. A method for flexural reinforcement of old wood beams with CFRP materials. Compos Part B Eng 2005;36(2):143–53. [35] Alam P, Ansell MP, Smedley S. Mechanical repair of timber beams fractured in flexure using bonded-in reinforcements. Compos Part B Eng 2009;40(2):95–108. [36] Jankowski LJ, Jasienko J, Nowak TP. Experimental assessment of CFRP reinforced wooden beams by 4-point bending tests and photoelastic coating technique. Mater Struct 2010;43:143–50. [37] EN 338. Structural timber. Strength classes; 2003. [38] MagmaTech Limited, RockBar. Technical data sheet (accessed 2011, 2013). [39] Raftery G. The use of advanced composite materials in the flexural reinforcement of low-grade glulam [Ph.D. thesis]. Civil Engineering, College of Engineering and Informatics, National University of Ireland Galway; 2010. [40] Raftery GM, Harte AM. Material characterisation of fast grown plantation spruce. In: Proceedings of the institute of civil engineers: struct and build, June, SB6, vol. 167; 2014. p. 380–386. [41] Raftery GM, Harte AM, Rodd PD. Bond quality at the FRP wood interface using conventional wood laminating adhesives. Int J Adhes Adhes 2009;29(2):101–10. [42] Raftery GM, Harte AM, Rodd PD. Bonding of FRP materials to wood using thin epoxy gluelines. Int J Adhes Adhes 2009;29(5):580–8. [43] EN 408. Timber structures. Structural timber and glued laminated timber. Determination of some physical and mechanical properties; 2003. [44] EN 386. Glued laminated timber. Performance requirements and minimum production requirements; 2001.