Department of Civil Engineering, Monash University, Melbourne, Australia 2

Transcription

1 Fourth Asia-Pacific Conference on FRP in Structures (APFIS 2013) December 2013, Melbourne, Australia 2013 International Institute for FRP in Construction JOINING PULTRUDED FRP TUBULAR COMPONENTS INTO SPACE LATTICED SHELL STRUCTURE JOINT DESIGN AND PERFORMANCE F.J. Luo 1, Y. Bai 2 and Y. Lu 1 1 Department of Civil Engineering, Monash University, Melbourne, Australia 2 Corresponding author. Department of Civil Engineering, Monash University, Melbourne, Australia. Yu.Bai@monash.edu ABSTRACT The development of space latticed shell structure using fibre reinforced polymer (FRP) tubular members is a promising practice wherein the key issue is to join different components together. In this study, a bolted sleeve joint in a staggered pattern was designed for joining square tubular pultruded FRP members into the proposed system inspired by Unistrut. The current study focuses on the joint performance under tensile loading and a total of four groups of specimen were prepared. Key parameters including bolt type (normal bolt or blind bolt), element dimensions (e.g. tube and bolt dimensions) and bolt group layouts were chosen and studied. During experiments, the failure modes, load-displacement and load-strain relationships were recorded to assess the joint performance. The test results demonstrate that a desirable ultimate load carrying capacity can be achieved by optimising element dimensions and bolt group layout when using normal bolt. Besides, when blind bolts are used, the compromise between easier fixing and joint performance must be considered. KEYWORDS Pultruded FRP profile, space latticed shell, bolted sleeve joint, joint performance INTRODUCTION Space latticed shell is a curved-surface structure in form of a network of structural elements which provides enclosure to a specific building or area, such as exhibition centre, sport stadium and transport terminal (Lan 2005) and is known for its great structural integrity, architectural aesthetic and advantages over some construction issues such as high degree of prefabrication, high efficiency of on-site installation and easy service installation (Chilton 2000). With the advent of high-performance composite materials, FRP materials have been introduced into civil engineering as alternatives to traditional construction materials with superior properties including lightweight, high-strength and anti-corrosion, making them a promising candidate for the development of space latticed shell structures. Nowadays, glass fibre reinforced polymer (GFRP) materials are more often used as structural members due to a lower cost in comparison to carbon fibre reinforced polymer (CFRP), although the latter exhibits higher strength and stiffness. For GFRP, however, there is an issue of low material s E-modulus making the serviceability design more critical. This may be somewhat compensated by an improvement in structural stiffness i.e. to assemble the structural members into latticed shell which is inherently rigid due to its curvature nature, triangular-based elements and sometimes the use of rigid connections (Lan 2005). In terms of manufacturing method, pultruded tubular FRP components would be the best choice over those from other manufacturing process (such as vacuum-assisted infusion) due to their modular nature which ensures highvolume cost-effective fabrication with consistent quality and requires less labour (Bank 2006). Such profiles are also characterised as high strength in pultrusion (axial) direction while a much lower strength and stiffness in transverse direction (Bai and Keller 2009). This implies a highly efficient use of such materials in latticed shell as the load transfer is predominately by axial forces (Lan 2005). As a general concern for FRP structures, one critical issue in developing FRP latticed shell is to join different components together whilst ensuring a reliable load transfer, higher structural integrity and efficiency. To implement such application, researchers have been seeking suitable joining methods since late 1970s and some relevant work are summarised as follows. The first and most popular approach is to bond FRP member with connector that matches a specific node. Typical examples were end fittings formed by crimping metallic sleeve, with the end portion threaded (Pickett et al. 1982) or flattened (Green and Phillips 1982) to fit different nodes. Besides, Hollaway and Baker (1984) developed a moulded glass filled nylon end cap end fitting and there were also variations of this design given in Hollaway (2010). The second method adopts mechanical fasteners to fix FRP tube on to the end fittings. For example, in Bai and Yang (2012), a single-bolt joint was used to assemble square tubular members to specially-designed multi-leg nodes for a proposed all-composite structure unit. Another example (Yonemaru et al. 1998) is a multi-row blind rivets jointed end fitting for Mero system (Lan 2005). Besides, there is also combination of adhesive and bolting technique, such as Haigo et al. (2003), where

2 the member was formed by inserting an inner tube with built-in bearing bolt into pultruded GFRP pipe and jointing them by multiple rows of rivets and adhesive. The fourth type of joint is featured by scaffolding in conjunction with a special erection process, where the members were stacked and jointed by swivel scaffolding elements, which allows them to go through a node continuously (Douthe et al. 2010). The fifth type of joint (Pfeil et al. 2009), adopted a dismountable and bearing type of connection between tubular FRP members and steel node, where the members and joints were held together by interior splice tubes and then pre-stressed to avoid them falling apart during service. Lastly, a 3D woven carbon fibre nodal joint was introduced in Stewart (2011), aimed at providing joining solution for all-composite space truss. Attempt was also made to produce universal node for all possible variations. Overall, the bonded and mechanical fastened joints are by far the most popular methods; however, neither is a perfect solution and studies mainly focus on circular member for Mero or Triodetric nodes which cover only a few cases of various successful proprietary systems as summarised in Lan (2005). For adhesive bonded joints, although with sufficient bonded area, a good mechanical performance can be achieved with less stress concentration and higher efficiency in contrast to bolted joints, there are issues such as the high cost of adhesive, difficulty in quality control for on-site assembly, degradation due to long-terms fatigue and environmental attacks and the likelihood of sudden and catastrophic failure with significant drop in load carrying capacity (Green and Phillips 1982; Duthinh 2000). Regarding mechanical fastening method, it is deemed to be more practical due to the ease of assembly, low maintenance and cost effectiveness (Matharu 2011); however, there are issues of severe stress concentration and the tendency of material being failed by shear under axial loading (Green and Phillips 1982; Hollaway and Baker 1984), which requires careful design. By far, only a few studies have been conducted for this method and other alternative fasteners also worth considering. The other methods, as above mentioned, are all highly promising (for example, type IV and V is not likely to fail by shear of which the longitudinal resistance is low for pultruded FRP) while in current stage they seem too sophisticated or casespecific to commercialize - this aspect also seems to be true for many of the bonded and mechanical fastened joints as proposed in previous studies. CONCEPTUAL DESIGN In this study, the proposed design is inspired by (Moduspan) Unistrut system (Chilton 2000) as shown in Fig.1 (a), where the steel member is replaced with modified structural components incorporating pultruded tubular FRP members with bended square tubular steel connector between the nodal plate and the FRP member [see Fig.1(b)]. The FRP members are jointed to steel connectors using bolt group forming a multi-row sleeve bolted connection. The primary objective is to develop a reliable solution to join square FRP tubular sections into a single-layer space latticed shell with all components mechanically fastened for practicability. It is also expected that, with adequate joint performance, this method will be also applicable to similar systems with other tubular sections. The current study focuses on the mechanical performance of the joint portion under axial tensile loading which has been identified as an important loading scenario for latticed shell structural members. (a) (b) Notes: 1 FRP structural member 2 Proposed bolted sleeve joint 3 Steel connectors 4 Nodal plate Figure 1. (a) Typical Unistrut System - after Chilton (2000); (b) proposed joining system in this study EXPERIMENTAL PROGRAM Materials The materials used included pultruded GFRP square tube (longitudinal material properties: tensile strength, 306±18 MPa, tensile modulus, 30.2±1.4 GPa and shear strength, 26.7±0.2 MPa; specifications: mm and mm) supplied by Nanjing Xingya FRP Co. Ltd., Grade 250 mild steel ( mm and mm SHS for the connectors and square solid for the plugs) from local supplier, Grade 8.8 High-tensile hex bolt (normal bolts) with steel zinc plated flat washers (1/4 1/2 18G and 5/16 5/8 18G) from local supplier (see Fig. 2) and high tensile zinc & yellow passivated blind bolt from Blind Bolt Australia (see Fig. 2). Slot Anchor (Normal Bolt) (Blind Bolt) Figure 2. Comparison of normal bolt and blind bolt

3 Specimens In this study, a total of four specimen sets were prepared with each set having three identical specimens. The specimens were designated according to the key parameters studied. Wherein, N or B, refers to the type of bolt i.e. normal bolt or blind bolt ; 1 or 2, refers to the number of bolt per row per side; M6 or M8 refers to the nominal bolt diameter; F refers to FRP ; and the last two numbers, 38 or 51 refer to the outer dimension of FRP tubes, corresponding to mm and mm, respectively. For compatibility, the steel tubes were selected to be mm for -F38 specimens, and for -F51 with overlap portion machined to mm. For design purpose, the proposed joint was considered as multi-bolted FRP-steel lap joint on an equivalent unfolding of tubular section. The key parameters such as bolt dimensions (nominal diameter and thread length), washer sizes, pitch distances, edge distances were determined according to the design requirement specified Section 5 of CNR-DT 205/2007 (2008) and Section 9 of AS4100 (1998). For each specimen, the number of bolt rows was kept as two while in a staggered pattern because small tubular sections do not allow a cross section having two bolts crossing orthogonally. The staggered distance (Δ) was kept minimal to avoid forming two separated rows whilst satisfying minimum geometry requirement from the above standards. Here, it is advantageous to have just two rows, as excessive number of rows does not significantly reduce the load transferred to the 1st bolt row (Hart-Smith 1980) and thus not cost-effective. Besides, all bolts were torqued to snug-tight condition to ensure elements solidly seated against each other. The details of specimen design are summarised in and Fig. 3 and Table 1. Row2 Row1 N1M6F38 N2M6F51 N1M8F51 B1M8F51 Figure 3. Schematic diagram of joint configuration (left) and specimens (right) Table 1. Details of specimen design Specimen Designations Specification of Bolts a d h e f e s S x S y Δ L f L s L m L N1M6F38 M6 60mm N2M6F51 M6 60mm N1M8F51 M8 75mm B1M8F51 M8 50mm Note: a d h bolt hole diameter; e f edge distance (FRP); e s edge distance (steel element); S pitch spacing. Experimental setup and procedure Specimens were tested in tension using 500kN-capacity Baldwin hydraulic-operated machine. The specimens were fixed by clamping the free ends of steel connectors between grip jaws to facilitate the transfer of tensile loading. To prevent the crushing of steel connector, the specimen ends were filled by fitting mild steel plugs according to ASTM A370 (see Fig. 3). The specimens were instrumented with strain gages on the rear side and one of the side faces to monitor the load-strain behaviour and strain distribution. In addition, Aramis photogrammetry was used to measure the strain distribution of the whole front side of the specimens. All specimens were tested beyond peak load at a loading rate of 2 mm/min. RESULTS AND DISCUSSION Failure modes In general, specimens were failed at ultimate load by shear-out from the end edge of FRP elements [Fig. 4(a)]. This is a typical failure mode for bolted connection on pultruded FRP materials loaded in the pultrusion direction and characterised by shear-out the part of laminate ahead of the bolt (Chamis 2013). Local failure as evidenced by crackling sound was observed in prior to the peak loading and the ultimate failure was associated with a sudden loud noise. The crack propagation, which was not easily observed by bare eyes, can be interpreted from Aramis strain distribution graph where the crack is found to be initiated at the end edge of FRP [Fig. 4(b)]. For

4 group N1M8F51, the specimens may be also failed by splitting on one side of the critical bolt hole pair [Fig. 4(c)] while the other side was still by shear-out. Beyond the ultimate load, the post-peak failure developed progressively at the critical end and the bolt movement was deemed to cease at the other end. Basically, shearout failure still remained predominant but associated with randomness in the bolt location at failure and the corresponding residual load. Another observation was the bolt flexure (or yielding) for B1M8F51 and N2M6F51 [Fig. 4(c-d)] initiated before the peak loadings. For B- specimens this would lead to bolt shear rupture [Fig. 4(e)] following by the failed bolts catapulted from their bolt holes, which was sudden and catastrophic. (a) (b) (c) (d) (e) Figure 4. (a) Typical shear-out failure from edge; (b) major strain distribution at ultimate load from Aramis; (c) typical splitting failure; (d) typical bolt flexure; (e) typical bolt shear rupture Load-displacement relationship The load-displacement relationship is summarised in Table 2 and Fig. 5 where only typical specimens were selected to illustrate. In common, all specimens undergo approximately linear elastic deformation up to ultimate load, followed by progressive loss of residual load carrying capacity. Also, shortly before the ultimate load was attained, there is a very small stiffness reduction for normal bolt specimens, which may attribute to the initiation of failure accompanying with crack propagation; for blind bolt specimens, the stiffness reduction is slightly higher, which is possibly due to the bolt yielding as aforementioned. Table 2. Summary of Test Results (left) Specimen Designations a Initial Stiffness (kn/mm) Ultimate Load (kn) N1M6F38-x 6.26 (±0.11) (±0.73) N2M6F51-x (±0.04) (±1.88) N1M8F51-x (±0.49) (±2.41) B1M8F51-x 6.97 (±0.06) (±3.87) Note: a initial stiffness is defined as the slope of the initial segment or the elastic part in Fig. 4 Figure 5. Comparison of typical load-displacement curves for each set of specimens Load-strain relationship and strain distribution The load-strain behaviour and strain distribution were investigated on a pair of adjacent sides covering the entire joint region and its vicinity using strain gages [Fig. 6(a)]. For illustration, strain results were selected only from N1M8F51. The strain data were grouped and compared according to their locations as shown in Fig.6 (b-f). Results are shown only up to the ultimate load as the subsequent strain readings were no longer reliable. On FRP elements [Fig.6 (b-e)] the regions ahead of bolt exhibited compressive strains while the rest were under tensile strains. For the ease of discussion, the net tension zone covering the 2 nd bolt row is called the inner (joint) region and the one covering 1 st row is call outer (joint) region. It was found that the inner region (G2/g2) exhibited similar tensile strains to the typical FRP region (G1/g1) outside the joint region, and the strain levels are about twice the outer section (G4/g4). In compressive regions, the strains ahead of the 1 st row of bolts (G5/g5) are generally higher than those of the 2 nd row (G3/g3) and the magnitude are also the highest amongst all gages, indicating that the 1 st bolt row tends to be the most critical one. Moreover, the gages in compressive regions may exhibit an initial zero or even tensile strain due to the delay of load pickup for certain bolt locations. In comparison, Gi and gi (i = 1, 2,.., 5) generally exhibited similar load-strain response as shown in Fig.6 (b-f). The variations identified from the above comparison may be because the load distribution is sensitive to bolt-hole clearance, tightening torque of bolt and the interfacial condition between elements for multi-bolted joints (Feo et al. 2012) which were hardly controlled in experiment. In addition, the strain levels are found independent to the bolt locations within a certain staggered row, indicating that the gap between adjacent staggered bolts are sufficiently close so that the load transfer would not differ into two distinct ways. Therefore, for design purpose, it would be reasonable to assume that the loading was equally shared by bolts in a certain staggered row, or

5 alternatively, bolt row in the proposed staggered pattern can be considered as in one single row in terms of load distribution. (a) (b) (c) (d) (e) (f) Figure 6. Load-strain behaviours of a typical specimen group (N1M8F51): (a) layout of strain gages; (b) ~ (f) load-strain relationships on different locations Effect of element dimension and bolt group layout From the comparison amongst normal bolt specimens N-, a higher ultimate load carrying capacity can be achieved by (1) using larger FRP member (N1M6F38 versus N1M8F51) or (2) increasing the number of bolts (N1M8F51 versus N2M6F51). This is not a surprise, as they all increase the shear-out strength that is directly proportional to edge distance and plate thickness according to the failure criteria as suggested in Cl of CNR-DT 205/2007 (2008). Meanwhile, the joint stiffness was also enhanced with the said strategies. Besides, since N2M6F51 provide a higher degree of structural redundancy than the other specimens with normal bolts due to the use of more fasteners, the post-peak failure process was more gradual and there were no sudden ladder-like load drop-offs (see Fig. 4). Effect of blind bolts Although blind bolt provide convenience for cavity fixing, it was found, from the comparison of B1M8F51 and N1M8F51, that the use of blind bolt would compromise the overall mechanical performance of joint (14.3% reduction in ultimate load and 30.9% in joint stiffness in reference to N1M8F51) due to the weakness of bolt itself i.e. the lack of shear resistance and bending stiffness due to the presence of slot which occupies nearly half of the cross section and separates the shear plane into a pair of circular segments (see Fig. 2), especially when the load was applied perpendicular to the slot direction. When yielding occurred on one bolt, it would not take further loading; instead, the load would transfer to the adjacent bolts making them more critical than in the reference scenario with normal bolt. CONCLUSIONS In this study, a bolted sleeve joint has been proposed for joining FRP members into space latticed shell structure. In current stage, four different specimens have been designed with varying bolt types, element dimensions and layouts, and the joint performance was examined under axial tension loading. The following conclusions can be drawn from this study. (1) The proposed joint exhibited linear elastic behaviour in load-displacement response

6 before failure and was predominately failed by shear-out from the end edge on FRP element at ultimate load. (2) Bolts in the same staggered row as proposed exhibited similar load-strain behaviour and therefore can be approximated as one single row with identical load distribution. (3) Using more bolts per row and larger section helps to achieve higher ultimate capacity due to the increasing shear-out resistance and the improvements were also found in the initial stiffness. (4) For the use of blind bolts, there is a compromise between the ease of cavity fixing and mechanical performance (14.3% reduction in ultimate load and 30.9% in joint stiffness by comparison of B1M8F51 and N1M8F51) due to the configuration of blind bolt itself. ACKNOWLEDGMENTS The authors would like to gratefully acknowledge the financial support provided by Australian Research Council Discovery Early career Research Award (ARC DECRA) and Monash University and the technical support from Mr. Long Goh and Mr. Mark Taylor (Civil Engineering Laboratory of Monash University). REFERENCES ASTM Standard A370-12a (2012), "Standard Test Methods and Definitions for Mechanical Testing of Steel Products" ASTM International, West Conshohocken, PA, Bai, Y. and Keller, T. (2009). "Shear failure of pultruded FRP composites under axial compression", Journal of Composites for Construction, 13(3), Bai, Y. and Yang, X. (2012). "Novel Joint for Assembly of All-Composite Space Truss Structures: Conceptual Design and Preliminary Study", Journal of Composites for Construction, 17(1), Bank, L. C. (2006). Composites for Construction: Structural Design with FRP Materials, John Wiley & Sons, Inc., Hoboken, New Jersey. Chamis, C. C. (2013). "Simplified Procedures for Designing Composite Bolted Joints", Journal of Reinforced Plastics and Composites, 9(6), Chilton, J. (2000). Space Grid Structures, Architectural Press, Butterworth-Heinemann, Oxford, U.K. CNR-DT 205/2007 (2008). Guide for the design and construction of structures made of thin FRP pultruded elements., National Research Council of Italy (CNR), Rome, Italy. Douthe, C., Caron, J. F. and Baverel, O. (2010). "Gridshell structures in glass fibre reinforced polymers", Construction and Building Materials, 24(9), Duthinh, D. (2000). Connections of Fiber-Reinforced Polymer (FRP) Structural Members: A Review of the State of the Art, NISTIR Feo, L., Marra, G. and Mosallam, A.S. (2012). "Stress analysis of multi-bolted joints for FRP pultruded composite structures" Composite Structures 94(12): Green, A. K. and Phillips, L. N. (1982). "Crimp-bonded end fittings for use on pultruded composite sections", Composites, 13(3), Haigo, H., Utsumi, Y., Kimura, K., Takahashi, K., Itohiya, G. and Tazawa, H. (2003). Development of space truss structure using glass fiber reinforced plastics, Adavanced materials for construction of bridges, buildings, and other structure III, ECI Digital Achives. Hart-Smith, L. J. (1980). Mechanically-Fastened Joints for Advanced Composites - Phenomenological Considerations and Simple Analyses, Fibrous Composites in Structural Design. E. Lenoe, D. Oplinger and J. Burke, Springer US: Hollaway, L. (2010). "Technical Papers: Polymers, Fibres, Composites and the Civil Engineering Environment: A Personal Experience", Advances in Structural Engineering, 13(5), Hollaway, L. and Baker, S. (1984). The development of nodal joints suitable for double-layer skeletal made systems made from fibre/matrix composites, Third International Conference on Space Structures : proceedings of the Third International Conference on Space Structures, H. Nooshin. University of Surrey, Guildford, UK, Elsevier. Lan, T. T. (2005). Space Frame Structures. Structural Engineering Handbook W. F. Chen, Boca Raton: CRC Press LLC. Pfeil, M. S., Teixeira, A.M.A.J. and Battista, R.C. (2009). "Experimental tests on GFRP truss modules for dismountable bridges", Composite Structures, 89(1), Pickett, A., Hollaway, L. and Phillips, L.N. (1982). "Analysis of a crimped and bonded joint for load bearing skeletal members", Composites, 13(3), Standard Australia, AS4100 (1998), Steel Structures Stewart, R. (2011). "Composites in construction advance in new directions", Reinforced Plastics, 55(5), Yonemaru, K., Fujisaki, T., Nakatsuji, T. and Sugizaki, K. (1998). "Structural properties of space truss structure with CFRP", Shimizu Technical Research Bulletin, (17),