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steel joints’ behaviour under subzero exposures
Accepted Manuscript Influence of fibres ’ stiffness on wet lay-up cfrp/steel joints ’ behaviour under subzero exposures Ahmed Al-Shawaf PII: DOI: Reference:
S1359-8368(14)00090-0 http://dx.doi.org/10.1016/j.compositesb.2014.02.013 JCOMB 2941
To appear in:
Composites: Part B
Received Date: Revised Date: Accepted Date:
13 October 2013 28 January 2014 14 February 2014
Please cite this article as: Al-Shawaf, A., Influence of fibres ’ stiffness on wet lay-up cfrp/steel joints ’ behaviour under subzero exposures, Composites: Part B (2014), doi: http://dx.doi.org/10.1016/j.compositesb.2014.02.013
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INFLUENCE OF FIBRES’ STIFFNESS ON WET LAY-UP CFRP/STEEL JOINTS’ BEHAVIOUR UNDER SUBZERO EXPOSURES Ahmed Al-Shawaf1 Department of Civil Engineering, Monash University, Victoria 3800, Australia [email protected] Department of Civil Engineering MONASH University 3800 Melbourne, Victoria, Australia T: +613 9879 7816 Fax: +613 9905 4944
Abstract The current paper characterizes and differentiates the bond behaviour of wet lay-up CFRP/steel double-strap joints fabricated with NM (normal modulus) and UHM (ultra high modulus) unidirectional CF (carbon fibres) plies. The influence of infrastructural subzero exposure on the bond attributes of these joints is also investigated. While environmental exposure is maintained at a particular subzero temperature, ranging from –40 to 20ºC, specimens representing these joints are tested in tension. Failure patterns, joints’ strength, and strain and LSS (lap-shear stress) distributions for the tested joints are extracted. Pertinent discussions and conclusions, related to the influence of CFRP moduli and subzero temperatures on the aforementioned parameters, are provided.
Keywords Cohesion failure; Load transfer; Normal modulus; Subzero exposures; Ultra high modulus; Wet lay-up
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Corresponding author
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Introduction Adhesively bonded CFRP/steel strap joints provide many advantages over the traditional steel-plate
bonding in retrofitting applications. These include remarkable strength and/or stiffness-to-weight ratio, faster field application due to its light weight, tailorable thermo-mechanical properties, relatively more durability as a material, higher chemical and corrosion resistance, and enhanced fatigue life [1]. In addition, it is a well-known fact that adhesively-bonded joints are strong in shear, but inevitably weak in peel. Therefore, double–strap joint configuration is chosen for the current work to ensure that the applied load is transferred mainly in shear and to minimize any direct or induced peel stresses which tend to detract from the joint’s capacity [2]. In CFRP retrofitting applications, CF reinforcement is commonly categorised into: (1) NM (Normal modulus) / high-strength, (2) HM (High modulus), and (3) UHM (Ultra high modulus). Ranges of CF elastic moduli for these categories generally fall into 230-240, 295-390, and 440-640 GPa, respectively [3]. However, the equivalent (i.e. overall) elastic moduli for the commonly-used preformed CFRP laminates varies from 150 to 450 GPa, respective to the lower and upper moduli limits of the aforementioned CF categories [4]. On the other hand, and for comparative CFRP layer’s thickness, wet/hand lay-up CFRP fabrication has considerably less range (e.g. 80 to 225 GPa [1]) due to practical limitations usually associated with this method. In this context, existing literature has a pronounced deficiency of pertinent wet lay-up strengthening publications compared to preformed methods [5-21]. Furthermore, there is even no universally accepted classification system, for the preformed methods, that researchers can refer to [22]. Therefore, it is proposed for the present work to ascribe wet lay-up CFRPs with moduli values below 100, 100-200, 200-400, and above 400 GPa as LM (Low Modulus), NM (Normal Modulus), HM (High Modulus), and UHM (Ultra High Modulus), respectively. It has been established that for any un-prestressed FRP strengthening material to be effective, its stiffness must be similar to the metallic structure being strengthened [3]. As such, NM CFRP is a cost-effective option for strengthening ‘carbon steel’ substrata as opposed to HM and UHM CFRP, since its strain and stress capacity can be mobilised after the substrate has yielded; and the strengthening material can be used considerably more effectively than with cast iron structures. On the contrary, cast iron is attributed with relatively low cracking strain, thus require FRP reinforcement with as high a modulus as possible to be effective in strengthening the section [3]. Consequently, HM and UHM CFRP would suit cast iron structures more efficiently. The majority of existing research on FRP strengthening for steel structures has mainly focused on preformed NM CFRP (i.e. CFRP modulus ≈ 100-200 GPa) [4, 22-26]. This research trend can be reasonably
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justified by the higher number, worldwide, of infrastructural carbon steel structures requiring CFRP strengthening compared to the other less-encountered type of metallic structures made from cast iron. Accordingly, more research should be directed towards HM and UHM CFRP external strengthening applications. The other under-investigated issue, which affects the infrastructural behaviour of CFRP/steel joints, is the environmental thermal exposures under which CFRP strengthening systems operate. Part of these exposures is subzero environmental exposures commonly encountered in infrastructural CFRP strengthening applications. These exposures are not expected to alter, tangibly, the thermo-mechanical attributes of reinforcing CF plies. However, the HM CFRP strengthening material is most likely to reveal different load transfer and failure mechanisms relative to LM and NM CFRP, under any specific subzero thermal exposure. This in turn would be reflected on the overall joint’s attributes, such as ultimate strength, and interfacial strain and LSS distributions [27, 28]. In the present investigation, it is intended to investigate the effect of utilising NM and UHM CF plies in fabricating wet lay-up CFRP/steel plate double-strap specimens on their bond behaviour under subzero thermal exposures. In order to avoid under-impregnated CF plies usually encountered in the wet lay-up method, precisely with the UHM type, a special fabrication rig for producing wet lay-up CFRP/steel specimens is adopted. The same steel adherend (i.e. carbon steel) shall be utilized for both CF types as the aim is to conduct bond behavioural comparison, with as minimum variables as possible. These specimens shall be tested in direct tension, under short-term sustained subzero environmental exposures ranging from –40 to 20°C. Assessments of failed specimens, at any subzero temperature, shall be based on their concluded failure patterns, capacities, longitudinal strain and LSS distribution within the joint’s interfacial adhesive layer.
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Testing program As mentioned earlier, double-strap joint with thin adherend (i.e. 4.85 mm steel plate) is adopted, as
this provides optimum configuration to ensure that the applied load is transferred mainly in mode II (i.e. shear), and to minimize any direct or induced peel stresses. Three CF plies were adopted as reinforcement for the wet lay-up CFRP laminate on each face of the double-strap joint. This number was concluded to be the optimum, in terms of load transfer efficiency, based on a previous investigation which adopted similar CF plies with wet layup method on steel plates [29]. The variables of the current specimens are: (1) elastic modulus of the CF plies
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(i.e. NM and UHM), and (2) subzero temperatures ranging from –40 to 20°C simulating thermal exposures encountered in outdoor CFRP retrofitting applications.
2.1
Wet lay-up joint’s constituents and processing The two unidirectional NM and UHM CF plies adopted for the current investigation are MBrace®CF
130 and 530, respectively. Both CF types were provided by Degussa Construction Chemicals Australia Pty Ltd. Their nominal thickness is 0.176 and 0.19 mm, respectively; and thermo-mechanical properties are quoted Table 1 according to the manufacturer’s technical data sheets [30, 31] and some literature standard-products data, as cited in the same table. Araldite®420 A/B is a commercial epoxide resin utilized as the matrix/adhesive for the present wet lay-up CFRP/steel double-strap joints, for both NM and UHM CF plies’ specimens. It was provided by Huntsman Advanced Materials. Its experimentally-extracted thermo-mechanical properties at current subzero environmental exposures are presented in Table 2.
Experimental procedures for characterising resin’s
parameters are detailed in Al-Shawaf [1]. Only the coefficient of thermal expansion was obtained from relevant adhesive’s literature (e.g. manufacturer data sheet). For the steel adherend, mild steel plates with 380 MPa yield strength, 227 GPa Young’s modulus, and nominal thickness of 4.85 mm are used in the current program. The steel plate was purchased from OneSteel Australia Limited. The wet lay-up CFRP/steel plate double-strap specimens were fabricated with a specially-designed steel rig shown in Figure 1. For detailed CFRP samples’ fabrication and rig’s design justification and advantages, readers are referred to Al-Shawaf and Zhao [33]. Each CFRP/steel sample was left for 7 days to cure under ambient exposure prior to water-jet cutting into 8 identical double-strap specimens, according to the predetermined dimensions.
2.2
Specimen setup, thermal conditioning, and testing ERSGs (electrical resistance strain gauges), compatible with plastic surfaces and with wide-range
limits of operational temperature from –60 to +250°C, were instrumented on the CFRP outer surface, as shown in Figure 2, to capture axial strain readings along the bondlength during load application. The application of the ERSGs and specimens’ setup, thermal conditioning, and testing procedures are detailed in Al-Shawaf and Zhao [33].
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2.3
Current CFRP/steel joints’ configurations Figure 3 (a and b) indicates the two configurations of the FRP joints adopted in this study. The details
of this figure are self-explanatory. As can be implied, the modulus for the wet lay-up CFRP strengthening material for the whole joint is twice that on each face of the double-strap configuration. Accordingly, the current NM CFRP/steel joint (i.e. Figure 3a) is anticipated to be more effective compared to the UHM CFRP/steel joint of Figure 3b, because its overall stiffness (i.e. ≈ 160 MPa) is considerably closer to that of the steel plate compared to the UHM CFRP/steel joint. Table 3 below and the following paragraph, facilitates the interpretation of the results, discussions, and analyses presented later. Specimens ID labels reflect the material configuration and exposure temperature for each of the experimental composite joints. The current subgroups are NA and HA. The first character denotes the modulus grade of the adopted CF plies, viz. (N) for NM CF (i.e. E = 240 GPa) and (H) for UHM CF (i.e. E = 640 GPa). The following (A) character acronyms the Araldite adhesive/matrix. Then, the environmental exposure temperature follows. Lastly, the replicated number of any particular specimen is indicated. For example, HA20,2 denotes the second specimen fabricated with UHM CF and conditioned at -20°C.
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Analyses of results Modes of failure, joints’ capacities, and interfacial longitudinal strain and LSS distributions, for all
tested CFRP/steel double-strap joints, are discussed in this section. Informative analyses on these parameters are properly addressed and justified. Comparative assessments of the effect of subzero temperatures on both types of specimens (i.e. NM and UHM CFRP/steel joints) are conducted, based on the abovementioned parameters.
3.1
Modes of failure Figure 4 is deemed useful in interpreting current failure modes. It depicts the likely modes of failure
encountered in FRP/ steel plate double-strap joints. As a prelude, any monotonically-loaded CFRP/steel joint upon failure displays one of two apparent failure modes’ types. These can be described as ‘single-mode’ and ‘mixed-mode’. In the former category, as the name implies, failure pattern is manifested wholly by any single mode of Figure 4. Practically, this case is rare; therefore the ‘mixed-mode’ is frequently prevailing. Nonetheless, few exceptions apply, such as bond failures at extremely elevated temperatures’ exposures close to the interfacial adhesive’s Tg (glass transition temperature),
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or if the interfacial adhesive/matrix is incompatible with steel adherends [1]. In mixed-mode failures, and with most commercially steel-compatible adhesives/matrices, actual bond capacity for well-configured CFRP/steel joints is closely related to the failure mode responsible for failure initiation (i.e. triggering failure). FEA (finite element analysis) method has proved to provide effective support in acquiring reasonable pre-experience required to determine the most likely triggering mode of failure [28]. Figure 5 assists the reader, at later discussions, in referring to the 4 potential sides (i.e. bondlengths) of the current joints’ failure locations.
3.1.1
NA-subgroup All specimens of this subgroup, at current subzero thermal exposures, exhibited mixed-mode failure
type of varying intensity, as indicated in Figure 6. The predominant failure mode was interlaminar failure in the CFRP laminate (i.e. Figure 4d). Primarily, this implies that the current fabrication rig was successful in providing quality adhesion and preventing premature interfacial debonding failure (i.e. Figure 4a). Moreover, ‘localized interfacial debonding’ (i.e. type (a) of Figure 4) were detected in the proximity of either/both ends of the instrumented bondlength. This pattern was found in all NA-subgroup specimens. It was found in Al-Shawaf [1], upon performing FEA of these joints, that their triggering failure pattern is ‘cohesion failure’ similar to Figure 4c and located close to either end, as well. As such, it is believed that current failures were triggered within the interfacial adhesive layer by a tensile cohesion failure close to either end of the bondlength creating scattered deep hackles, which were then propagated and eventually linked up at the adhesive/steel interface before the CF/adhesive interface due the continuum nature of the steel surface, as opposed to the filamentary nature of the carbon fibres. Hence, those localized interfacial debonding patches were observed, experimentally. For more details on this failure mechanism, readers are referred to Hart-Smith [2] / Al-Shawaf and Zhao [33]. The latter failure pattern for the NA-subgroup specimens has not been influenced by the current range of subzero exposure temperatures (i.e. –40, –20, and 0°C). The only observed difference was the increased brittleness and violent nature of the failure process as the conditioning exposure temperature of the specimens decreased.
3.1.2
HA-subgroup Tested specimens belonging to this subgroup are portrayed in Figure 7.
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The apparent predominant failure mode for the HA-subgroup specimens is CFRP rupture at the joint (refer to Figure 4e). Also, very small areas (about 2% of the overall bond area for the four bondlengths) of ‘localized interfacial debonding’ (i.e. type (a) of Figure 4), close to the 2 mm gap were visually detected, especially for the 20°C-exposure specimens. This failure pattern can be attributed as ‘fibre-controlled’, as opposed to the ‘adhesive-controlled’ failure pattern found in the NA-subgroup specimens. The dissimilarity of the current failure pattern from the NA-subgroup is related to the difference in joints’ load transfer mechanism, and thus interfacial adhesive stress distribution. This different load transfer mechanism is basically a direct consequence of lower tensile strength and ultimate elongation values (i.e. 2070 MPa and 0.4%, respectively) of the UHM CF plies compared with their respective NM CF plies’ values of the NA-subgroup, viz. 3800 MPa and 1.55%, respectively. It was found, via FEA for the HA-subgroup joints that although 1st CF ply’s rupture at the joint initiated failure, its propagation would not have been facilitated had the strain energy level of the neighbouring interfacial adhesive layer been relatively low. Moreover, FEA in the vicinity of the mid-gap showed that HA–40 specimens had simultaneously attained ultimate capacities for the 1st CF ply and the interfacial adhesive/matrix material [1]. Accordingly, it can be concluded that the common instigator of the HA-subgroup joints’ triggering failure is related, either directly or indirectly, to the interfacial adhesive layer, viz. its stress status. As with the NA-subgroup, current subzero thermal exposures had almost negligible effect on the predominant failure mode trend.
3.2
Ultimate joints’ capacity Table 4 presents the ultimate loads for current specimens. In order to be consistent in concluding the
short-term effect of environmental exposures on joint capacity, the triggering failure pattern should be identical for all specimens involved in any comparison representing the whole temperature spectrum. From Subsection 03.1, this prerequisite is found applicable for the present work.
3.2.1
NA-subgroup Based on the preceding paragraph, all specimens of the NA-subgroup developed identical failure
patterns, both triggering and predominant. As such, noting their corresponding average ultimate loads at current subzero and ambient exposures in Table 4, no special trend was inferred from increasing the short-term exposure temperature from –40°C to 20°C. However, by comparing the relevant joints’ capacities with those
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extracted from corresponding simulated FE models [28], close validation was concluded for the FE models of the NA-subgroup specimens with the higher experimental capacities amongst the two replicates of each exposure (i.e. NA–40,1, NA–20,1, NA+0,1, and NA+20,2). Consequently, and taking into consideration the expected experimental results’ scatter between replicated specimens, the ones with higher bond capacities can be assumed as representative of the accurate behaviour at each environmental exposure. The ultimate capacities of those specimens suggest that there is a reduction of about 4.5% for each 20°C temperature increase within the current subzero temperatures’ range , which is generally in line with few previous investigations [5, 34].
3.2.2
HA-subgroup The generic ultimate load capacity trend for specimens of this subgroup is characterized by lower
values than the NA-subgroup. Apparently, the direct reason for this observation is the considerably stiffer and lower tensile strength of the UHM CF plies of the HA-subgroup specimens as compared to the NA-subgroup counterparts. The latter trend is particularly evident at the subzero temperatures. From Table 4, it can be concluded that there is no tangible difference between the HA-subgroup average joints’ capacities at –40 and –20°C exposures; however they increased by 17% and 5% at 0 and +20°C exposures, respectively. It is believed that the abovementioned variations are related to the different contributions of the adhesive/ matrix in the load transfer mechanism. More precisely, the substantial 35% increment in the adhesive’s ‘elongation at failure’ as the conditioning temperature increased from –20 to 0°C is a reasonable justification of the aforementioned 17% joint’s capacity increment. This 35% extra elongation capacity simply enhanced the interfacial adhesive’s ability of accommodating more shear lag deformations before failure. Consequently, the average joint’s LSS, and thus ultimate capacity, have increased. The same scenario is proportionately applicable as the thermal exposure increased from 0 to 20°C. This behaviour clearly highlights the role played by the rheological properties of the adhesive/matrix in determining the individual load transfer mechanism, at each subzero environmental exposure, and its key influence in shaping the joint’s failure pattern and ultimate capacity.
3.3
Axial strain distributions Strain readings along the instrumented bond length were captured via standard ERSGs (number 1 to 7
as indicated in Figure 2). The total and gauge lengths for the current ERSGs are 13 and 6 mm, respectively. It has been demonstrated that, with standard ERSGs, some experimental / practical strain-measuring constraints have to be considered in order to meaningfully assess the accuracy of strain readings, and thus LSS. In this
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study, these strain-measuring constraints exist at bondlength ends (i.e. locations of SGs 1 and 7). Accurate and representative strain readings at these locations are paramount due to their close relevance to joint’s triggering failure proximity. Close to x = 0, strain/stress exhibits sharp peak away from directly above the 2-mm gap (Figure 2) due to some limited bending of the CFRP reinforcement at this point. This bending, which is undetectable by SG 1 due to its 6-mm gauge length (i.e. it is actually overlapping the 2-mm gap), is believed to cause a compressive stress/strain at the surface (i.e. outer side) of the CFRP layer within this narrow proximity [28]. At the other end (i.e. exactly at x = 100 mm), it is practically impossible to obtain strain readings at the CFRP’s edge with traditional strain gauging techniques. For these reasons, and more broadly, the complexity of the stress distributions in bonded joints; their theoretical analyses is very difficult to verify experimentally, which renders FEA methods as the optimum recommended option for accurate behavioural predictions within the interfacial adhesive layer [35]. Apart from bondlength ends, accurate experimental strain distributions were obtained for both NA and HA-subgroups, from x = 16 to 96 mm. These where validated via FEM predictions of the current experimental joints [1]. It has been assumed that the inherent existence of shear lag deformations within CFRP laminate’s thickness is ignored due to the relatively stiff response of most structural-grade adhesives at subzero to ambient thermal exposures [1]. In addition, it is assumed that bulk longitudinal and shear deformations in the adherend (i.e. steel plate) can also be neglected due to their small values relative to those of the wet lay-up CFRP laminate and adhesive layer.
3.3.1
NA-subgroup Figure 8 depicts the axial strain readings, for one replicate from each thermal exposure of the NA-
subgroup specimens, plotted with distance from the joint, for 10% successive increments of ultimate failure load. The general pattern in the strain distributions for subzero and ambient exposures is, as expected, ultimate strain values at the joint (i.e. x = 0) followed by gradual reductions towards the other end of the bondlength at all load levels (i.e. from 10% Pult. to Pult.); with the rate of strain reduction decreasing sharply from x = 0 to x = 32 mm, after which almost uniform distributions are displayed up to x = 80 mm, then approaches zero at x = 96 mm (i.e. location of SG7). This trend is consistent with the traditional stiffness-imbalanced and thermally-mismatched doublelap joint model [36]. The latter model establishes that, for a double lap-shear joint, whenever the net stiffness at one end of a joint differs from that at the other end, the adhesive shear strain distribution is rendered non-
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symmetric. The second cause of strain non-symmetry is the thermal mismatch between adherends, according to the same model. In case of differences in coefficients of thermal expansion between dissimilar adherends, more severe strains are associated with the end where the lower coefficient of thermal expansion’s adherend extends. From Al-Shawaf [1], the coefficients of thermal expansion for the NA-subgroup CFRP laminate are 2.408, 2.042, 1.620, and 1.521x10–6/°C at –40, –20, 0, and 20°C thermal exposures, respectively; and for steel is 11x10-6/°C. Therefore, thermal mismatch effect has actually added up shear strains close to x = 0 end of the joint. Comparison of the 4 specimens in Figure 8 reflects the anticipated effect of increasing thermal exposure from –40°C to 20°C. This effect is manifested in slightly increasing both the uniformity of strain distributions along the bondlength and the ultimate strain values at the joint. The latter trends are direct consequences of decreasing embrittlement (i.e. increasing ductility) with temperature for joints with ductile adhesives, such as the current one.
3.3.2
HA-subgroup The axial strain distributions for selected HA-subgroup specimens, representing all currently-
experimented thermal exposures, are presented in Figure 9. In terms of the general trend, it resembles the previously-discussed one in that the highest strain is always at x = 0, followed by gradual declination in strain values towards the other end of the instrumented bondlength (i.e. x = 100 mm). Two outstanding attributes can be easily concluded from the comparison of Figures 9 and 8. The first one is the negligible difference amongst the ultimate strain values for the four specimens of the HA-subgroup compared to the respective tangible strain increments with conditioning temperatures for the NA-subgroup. This aspect supports the ‘fibre-controlled’ load transfer mechanism of the former subgroup as opposed to the ‘adhesive-controlled’/ ‘adhesive dependant’ for the latter. The second is the approximately 78% drop in the ultimate strain values at the joint (i.e. x = 0) exhibited by the HA-subgroup specimens relative to the NAsubgroup ones which clearly reflects the considerable stiffness’s difference between both CFRPs.
3.4
Lap-shear stress distributions The average LSS between two adjacent ERSGs was calculated by dividing the force difference by the
bond area between both gauges (per unit width of the bond), according to the following equation:
τ = E frp . (ε f i +1 − ε f i ). t frp / Δl
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(1)
where Efrp and tfrp are the CFRP elastic modulus and thickness, respectively; εfi+1 and εfi are the CFRP axial strains (i.e. strain gauge readings) at two consecutive locations Δℓ distance apart (i.e. 16 mm as indicated in Figure 2). Stresses calculated via Equation (1) actually represent the average shear stresses at 16 mm intervals along bondlength, thus are smaller than actual values in the composite specimen. A dominant trend of very good agreement was detected between experimental LSS distributions and their predicted FE counterparts, from x = 24 to 88 mm along instrumented bondlength, for both subgroups and at all load levels [1]. Locations closer to both bondlength ends revealed some divergence from the latter trend due to practical reasons explained earlier (i.e. in the prelude to Section 3.33.3).
3.4.1
NA-subgroup LSS distributions along instrumented bondlength for one replicate of the NA-subgroup specimens are
presented in Figure 10. The general trend was successfully verified against existing analytical model/s in AlShawaf [1]. The established trend of LSS distribution in a balanced double-lap joint (i.e. Eiti = 2Eoto) is the nonuniformity of the load transfer throughout the adhesive bondlength, which peaks at each end of the overlap and is symmetrical about a perpendicular line passing through x = ℓ/2 of the bondlength [2]. In the current investigation, due to adherend stiffness imbalance and the thermal mismatch of all double-strap joint specimens, the aforementioned symmetry is not the case. Instead, the prevailing trend for the NA-subgroup specimens under all environmental exposures is the non-symmetry in LSS distribution, with peak values occurring close to x = 0 end. From Figure 10, the embrittlement effect at subzero temperatures is clearly dominating the LSS behaviour of specimens NA–40,2, NA–20,1, and NA+0,1 as compared to specimen NA+20,2. It can also be detected that the –40, 0, and 20°C exposure specimens have followed similar trend close to x = 0 end of the bondlength, and for almost all load levels where the adhesive behaviour is linear. The only exception to this trend occurred with both specimens NA–20,1 and 2. This distinct LSS trend, at –20°C exposure, can be traced back from the corresponding axial strain distribution in Figure 8. Apparently, this could be misinterpreted as adhesive non-linearity, however close examination of specimen’s NA–20,1 LSS distribution (refer to Figure 10) reveals comparable stress values at both x = 8 and 24 mm locations, even at low-load levels where adhesive usually behaves linearly. The likely explanation to this phenomenon is believed to be due to different thermal residual stress’s response for the CFRP laminate at each conditioning temperature. These residual thermal stresses add up to the mechanical stresses to shape up the final bond LSS distribution.
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3.4.2
HA-subgroup Figure 11 portrays the interfacial LSS distributions along the instrumented bondlengths of four
representative specimens from HA-subgroup. It is believed that LSS distributions are of secondary significance in CFRP joints where load transfer and failure mechanisms are driven by the reinforcing fibres (i.e. fibrecontrolled), as the case for HA-subgroup specimens. As such, the following discussion is merely to stress the behavioural difference of LSS distributions between both NA and HA-subgroups. The same characteristic LSS distribution trend of peak values close to bondlength ends is applicable for the specimens of this subgroup. However, the key variation reflected on the HA-subgroup’s specimens relative to the NA subgroup’s is the conversion of the LSS peaks from ‘x = 0’ end, for all specimens of the latter subgroup, to ‘x = 100 mm’ end for the corresponding specimens of the former subgroup. This variation was apparently by virtue of the difference occurring to the combined effect of adherends’ stiffness imbalance and thermal mismatch. The other useful implication extracted from Figure 11 is related to the HA-subgroup specimens’ strengths compared to the NA-subgroup ones. From both Figure 10 and Figure 11, the average LSS of specimens belonging to the former is apparently higher than that for correspondent specimens of the latter. This finding is useful in predicting/justifying the higher joints values, presented earlier, for the NA-subgroup specimens compared to their counterparts from the HA-subgroup. In terms of the effect of subzero exposures, no clear trend can be implied on the LSS distributions of the HA-subgroup specimens. The ‘fibre-controlled’ behaviour of the current specimens is anticipated to, reasonably, justify this trend.
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Conclusions The prime conclusion inferred from the current study is the substantial role of the adhesive / matrix
for wet lay-up CFRP/steel joints reinforced with NM CF plies in the load transfer process, as opposed to the predominance of the CF plies’ role for similar joints, however, reinforced with UHM CF plies. More specific parameter-related conclusions for the wet lay-up fabrication method are drawn in the following subsections:
4.1
Failure modes
• Current subzero environmental exposures have almost negligible effect on the individual predominant failure mode for double-strap NM CFRP/steel joints and UHM CFRP/steel joints.
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• Double-strap NM CFRP/steel joints manifest ‘adhesive-controlled’ characteristic failure mode; while UHM CFRP/steel joints reveal clear ‘fibre-controlled’ mode. • For ductile interfacial adhesives, the likely triggering failure pattern for double-strap NM CFRP/steel joints is cohesion failure in the adhesive layer close to either ends of the bondlength, and CFRP rupture for UHM CFRP/steel joints.
4.2
Bond strengths
• Double-strap CFRP/steel joints manifesting ‘Adhesive-controlled’ behaviour usually attain higher capacities than identical joints disclosing ‘fibre-controlled’ behaviour. • Capacities for UHM CFRP/steel joints are more susceptible to variation in subzero thermal exposures within current range than NM CFRP/steel joints fabricated with identical ductile interfacial adhesive.
4.3
Axial strain distributions
• For CFRP/steel joints fabricated with identical ductile adhesive, ultimate strain values are inversely proportional with CF plies’ stiffness. • Axial strain distributions for NM CFRP/steel joints are more vulnerable to thermal variation within current subzero exposures than UHM CFRP/steel joints. • Standard ERSGs are incapable of capturing accurate strain values at bondlength ends where load transfer is highest. Instead, innovative strain-measuring techniques should be researched, such as FOS (fibre optic sensors), advanced non-contact optical and scanning electron microscopy (SEM) methods.
4.4
LSS distributions
• Measuring interfacial LSS distributions experimentally is more useful in predicting failure of CFRP/steel joints attributed with ‘adhesive-controlled’ behaviour; as opposed to ‘fibre-controlled’ behaviour’s joints. • The generic LSS distribution trend conforms to the traditional model of linear distribution with its characteristic peak values close to both ends of the bondlength at loads below the triggering failure one. • Ultimate LSS’s location is function of the CFRP layer’s stiffness. It is close to x = 0 mm in joints with low CFRP stiffness, as opposed to x = ℓ for joints with high CFRP stiffness. • As with the strain distribution trends, LSS distributions for NM CFRP/steel joints are more susceptible to thermal variations, within current subzero exposures, than UHM CFRP/steel joints.
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Acknowledgements This work is sponsored by an Australian Research Council (ARC) Discovery Grant. The kind
assistance of Jim Mitchell and Silvio Mattievich in setting up and operating the environmental chamber and the Instron testing machine at the Department of Materials Engineering / Monash University is acknowledged by the author. He also wishes to thank Kevin Nievaart for his contribution in facilitating the progress of the experimental part at both the Civil and Instron laboratories. Finally, the great help of all staff at the Civil Laboratory / Monash University is appreciated in executing the experimental program.
References 1. Al−Shawaf A. Characterization of bonding behaviour between wet lay-up carbon fibre reinforced polymer and steel plates in double-strap joints under extreme environmental temperatures. Doctorate PhD thesis, Monash University, Melbourne, Australia, 2010. 2. Hart−Smith LJ. Bolted and bonded Joints. In: Miracle DB, Donaldson SL, editors. Composites. Ohio, USA: ASM International, 2001. p.271-89. 3. Cadei JMC, Stratford TJ, Hollaway LC, Ducket WG. Strengthening metallic structures using externally bonded fibre-reinforced polymers C595. London: CIRIA, 2004. 4. Zhao XL, Zhang L. State of the art review on FRP strengthened steel structures. Eng Struct 2007;29(8): 180823. 5. El−Hacha R, Green MF, Wight G. Strengthening of concrete members with prestressed FRP sheets: Behaviour at room and low temperature. London, Ont., Canada: Canadian Society for Civil Engineering, Montreal, H3H 2R9, Canada, 2000:187. 6. di Tommaso A, Neubauer U, Pantuso A, Rostasy FS. Behavior of adhesively bonded concrete-CFRP joints at low and high temperatures. Mech Compos Mater 2001;37(4): 327-38. 7. Bisby LA, Green MF. Resistance to freezing and thawing of fiber-reinforced polymer-concrete bond. ACI Struct J 2002;99(2): 215-23. 8. Colombi P, Fanesi E, Fava G, Poggi C. Durability of steel elements strengthened by FRP plates subject to mechanical and environmental loads. CCC2005 : Third International Conference Composites in Construction. Lyon, France, 2005. 9. Green MF. FRP repair of concrete structures: performance in cold regions. International journal of materials and product technology 2007;28(1-2): 160.
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10. Karbhari VM, Rivera J, Dutta PK. Effect of short-term freeze-thaw cycling on composite confined concrete. J Compos Constr 2000;4(4): 191-97. 11. Abanilla MA, Karbhari VM, Li Y. Interlaminar and intralaminar durability characterization of wet layup carbon/epoxy used in external strengthening. Compos Part B-Eng 2006;37(7-8): 650-61. 12. Agarwal A, Foster SJ, Hamed E, Ng TS. Influence of freeze–thaw cycling on the bond strength of steel–FRP lap joints. Compos Part B-Eng 2014;60(0): 178-85. 13. Al−Emrani M, Kliger R. Analysis of interfacial shear stresses in beams strengthened with bonded prestressed laminates. Compos Part B-Eng 2006;37(4-5): 265-72. 14. Almusallam TH, Al-Salloum YA. Ultimate strength prediction for RC beams externally strengthened by composite materials. Compos Part B-Eng 2001;32(7): 609-19. 15. Camata G, Spacone E, Zarnic R. Experimental and nonlinear finite element studies of RC beams strengthened with FRP plates. Compos Part B-Eng 2007;38(2): 277-88. 16. Colombi P, Poggi C. An experimental, analytical and numerical study of the static behavior of steel beams reinforced by pultruded CFRP strips. Compos Part B-Eng 2006;37(1): 64-73. 17. Czaderski C, Rabinovitch O. Structural behavior and inter-layer displacements in CFRP plated steel beams – Optical measurements, analysis, and comparative verification. Compos Part B-Eng 2010;41(4): 276-86. 18. Keller T, Vallée T. Adhesively bonded lap joints from pultruded GFRP profiles. Part I: stress–strain analysis and failure modes. Compos Part B-Eng 2005;36(4): 331-40. 19. Kim YJ, Harries KA. Behavior of tee-section bracing members retrofitted with CFRP strips subjected to axial compression. Compos Part B-Eng 2011;42(4): 789-800. 20. Parvin A, Granata P. Investigation on the effects of fiber composites at concrete joints. Compos Part B-Eng 2000;31(6–7): 499-509. 21. Yu T, Fernando D, Teng JG, Zhao XL. Experimental study on CFRP-to-steel bonded interfaces. Compos Part B-Eng 2012;43(5): 2279-89. 22. Peiris NA. Steel beams strengthened with ultra high modulus CFRP laminates. Doctorate PhD thesis, University of Kentucky, Lexington, 2011. 23. Zhao XL. Recent developments in FRP strengthening of metallic structures. 1st Middle East conference on smart monitioring, assessment and rehabilitation of civil structures. Dubai, UAE, 2011. 24. Yu Y, Chiew SP, Lee CK. Bond failure of steel beams strengthened with FRP laminates – Part 2: Verification. Compos Part B-Eng 2011;42(5): 1122-34.
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25. Sen R, Liby L, Mullins G. Strengthening steel bridge sections using CFRP laminates. Compos Part B-Eng 2001;32(4): 309-22. 26. Linghoff D, Al-Emrani M, Kliger R. Performance of steel beams strengthened with CFRP laminate – Part 1: Laboratory tests. Compos Part B-Eng 2010;41(7): 509-15. 27. Adams RD. Adhesive bonding : science, technology and applications. Cambridge: Woodhead Publishing Ltd, 2005:xvi, 543. 28. Al−Shawaf A. Modelling wet lay-up CFRP–steel bond failures at extreme temperatures using stress-based approach. Int J Adhes Adhes 2011;31(6): 416-28. 29. Fawzia S, Al−Mahaidi R, Zhao XL, Rizkalla SH. Strengthening of circular hollow steel tubular sections using high modulus CFRP sheets. Constr Build Mater 2007;21(4): 839-45. 30. BASF. MBrace® CF130 technical data sheet www.basf-cc-la.com, 2006. 31. BASF. MBrace® Fibre technical data sheet www.basf-cc.com.au, 2006. 32. BP−Amoco. Thornel carbon fibres-http://www.cytec.com/engineered-materials/downloads/ThornelTP.pdf. Available from URL: http://www.cytec.com/engineered-materials/downloads/ThornelTP.pdf. 33. Al-Shawaf A, Zhao XL. Adhesive rheology impact on wet lay-up CFRP/steel joints’ behaviour under infrastructural subzero exposures. Compos Part B-Eng 2013;47(0): 207-19. 34. Karbhari VM, Shulley SB. Use of composites for rehabilitation of steel structures - determination of bond durability. J Mater Civil Eng 1995;7(4): 239-45. 35. Harris JA, Adams RD. Strength prediction of bonded single lap joints by non-linear finite element methods. Int J Adhes Adhes 1984;4(2): 65-78. 36. Hart−Smith LJ. Adhesive-bonded double-lap joints. Hampton, Virginia, USA: NASA Langley Research Centre, 1973:106.
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6
Figure Captions
Figure 1 CFRP/steel double-strap joint fabrication rig, (a) schematic section [not-to-scale], and (b) photo of assembled rig Figure 2 Geometry and instrumentation of experimental program’s specimens (Not-to-scale) Figure 3 The two currently-adopted wet lay-up CFRP/steel double-strap joints’ configurations (not-to-scale) Figure 4 Possible modes of failure for double-strap configuration of FRP/steel adherend joint Figure 5 Schematic key diagram for designating the four potential failure sides for the current double-strap joint configuration Figure 6 Failure patterns for the NA-subgroup specimens Figure 7 Failure patterns for the HA-subgroup specimens Figure 8 Axial strain distributions along instrumented bondlength for NA-subgroup specimens Figure 9 Axial strain distributions along instrumented bondlength for HA-subgroup specimens Figure 10 LSS distributions along instrumented bondlength for NA-subgroup specimens Figure 11 LSS distributions along instrumented bondlength for HA-subgroup specimens
17
7
Tables
Table 1 Nominal properties for MBrace®CF 130 and 530 carbon fibre plies CF category MBrace®CF 130 MBrace®CF 530 * [19]
Long. Specific thermal heat conductivity (j/Kg.°C) (j/s/mm/°C)
Modulus of elasticity (GPa)
Long. strength (MPa)
Ultimate elongation (%)
Shear strength (MPa)
Long. Poisson’s ratio
Long. CTE (x 10-6/°C)
240
3800
1.55
117.2 *
0.27 *
-0.38
9.38
711 *
640
2070 *
0.4
50 *
0.245 *
-0.83
68.7
711 *
18
Table 2 Typical thermo-mechanical properties for Araldite®420 A/B epoxide resin Thermal Tensile exposure modulus (MPa) (°C) –40 –20 0 20
2974 2575 2118 2012
Poisson’s ratio
Tensile strength (MPa)
Elongation at failure (%)
0.364 0.346 0.333 0.366
33.26 43.90 40.17 32.60
1.24 1.99 2.68 2.85
Coefficient of Glass Thermal thermal conductivity transition expansion [W/m.°C] temp. (Tg) [°C] [x10-6/°C] 0.1431
19
41.66
110
Specific heat (j/g.°C) 0.8985 0.9679 1.0207 1.1328
Table 3 Current CFRP/steel double-strap specimens IDs Adopted CF category ®
MBrace CF 130 MBrace®CF 530
–40 NA–40,1 NA–40,2 HA–40,1 HA–40,2
Environmental exposure (°C) –20 0 NA–20,1 NA+0,1 NA–20,2 NA+0,2 HA–40,1 HA+0,1 HA–40,2 HA+0,1
20
20 NA+20,1 NA+20,2 HA+20,1 HA+20,2
Table 4 Ultimate loads of the current CFRP/steel double-strap specimens NA-subgroup Joint Specimen capacity (kN) NA–40,1 94.8 NA–40,2 75.6 NA–20,1 89.3 NA–20,2 77.3 NA+0,1 88.9 NA+0,2 76.5 NA+20,1 82.5 NA+20,2 88.7
Average capacity (kN) 85.2 83.3 82.7 85.6
HA-subgroup Joint Average Specimen capacity capacity (kN) (kN) HA–40,1 58.9 58.5 HA–40,2 58.0 HA–20,1 69.1 58.1 HA–20,2 47.1 HA+0,1 70.5 68.1 HA+0,2 65.6 HA+20,1 78.8 71.7 HA+20,2 64.5
21
8
Figures’ legends
•
Figure 8 Legend (identical for the 4 specimens)
•
Figure 9 Legend (identical for the 4 specimens)
•
Figure 10 Legend (identical for 4 specimens)
•
Figure 11 Legend (identical for 4 specimens)
22
Figure
Two screw-driven press plates
Press supporting frame
Steel angle
Wet Lay-up CFRP laminates (4.85 mm) Two sample steel plates
Steel angle
(2mm) Shims plates
(a)
(b)
Figure
4.85 mm
2 Wet lay-up CFRP layers
S.G.9
2 mm 2 mm
S.G.8
2-mm gap
steel plate
steel plate
S.G.1 S.G.2 S.G.3 S.G.4 S.G.5 S.G.6 S.G.7
Side view Total specimen length = 302 mm 6 @ 16.0 mm
Width = 50 mm S.G.1 S.G.2 S.G.3 S.G.4 S.G.5 S.G.6 S.G.7
steel plate
S.G.8
Bottom face view 100 mm
100 mm S.G.9
wet lay-up CFRP length = 202mm
Top face view
steel plate
Figure
LM CFRP laminate Ecfrp ≈ 80 GPa
4.85 mm
2 mm
NM CF-3 plies Eply = 240 GPa
2 steel plates
Identical bottom wet LM CFRP laminate (a) NM CFRP/steel double-strap joint (i.e. total CFRP modulus ≈ 160 GPa)
HM CFRP laminate Ecfrp ≈ 225 GPa 2 mm
UHM CF-3 plies Eply = 640 GPa
4.85 mm
Identical bottom wet HM CFRP laminate (b) UHM CFRP/steel double-strap joint (i.e. total CFRP modulus ≈ 450 GPa)
Figure
FRP laminate
Adhesive
(a) Substrate-adhesive debonding
(b) Adhesive-FRP debonding
(c) Cohesion failure
(d) Interlaminar failure in the FRP laminate
(e) FRP rupture at the joint
Figure
Side C
Side A
Side D
Side B
Figure
Failure sides: A + D Failure pattern: localized type (a) + type (d)
a
Failure sides: A + B Failure pattern: localized type (a) + type (d) a
Failure sides: A + D Failure pattern: localized type (a) + type (d) a
Failure sides: A + B Failure pattern: localized type (a) + type (d) a
Failure sides: A + D Failure pattern: localized type (a) + type (d) a
Failure sides: A + B Failure pattern: localized type (a) + type (d) a
Failure sides: A + B Failure pattern: localized type (a) + type (d) a
Failure sides: A + B Failure pattern: localized type (a) + type (d) a
a
Refer to Figure 5 for interpretation of failure mode types.
Figure
Failure sides: CFRP rupture (B/D) + C Failure pattern: type (e) + type (d) a
Failure sides: Total CFRP rupture Failure pattern: type (e) a
Failure sides: Total CFRP rupture Failure pattern: type (e) a
Failure sides: Total CFRP rupture Failure pattern: type (e) a
Failure sides: Total CFRP rupture Failure pattern: type (e) a
Failure sides: Total CFRP rupture Failure pattern: type (e) a
Failure sides: Total CFRP rupture Failure pattern: type (e) a
Failure sides: Total CFRP rupture Failure pattern: type (e) a
a
Refer to Figure 5 for interpretation of failure mode types.
Figure
0.2 Pult 0.7 Pult
0.3 Pult 0.8 Pult
0.4 Pult 0.9 Pult
0.1 Pult 0.6 Pult
0.5 Pult Pult
12000
12000
10000
10000
Axial strain (με)
Axial strain (με)
0.1 Pult 0.6 Pult
8000 6000 4000 2000
0.3 Pult 0.8 Pult
0.4 Pult 0.9 Pult
0.5 Pult Pult
8000 6000 4000 2000 0
0
0
16
32
48
64
80
(x) distance from joint (mm) NA–40,2 0.1 Pult 0.2 Pult 0.3 Pult 0.4 Pult 0.6 Pult
0.7 Pult
0.8 Pult
0.9 Pult
0
96
0.1 Pult 0.6 Pult
0.5 Pult Pult
16
32
48
64
80
96
(x) distance from joint (mm) 0.2 PultNA–20,1 0.3 Pult 0.4 Pult 0.5 Pult 0.7 Pult
0.8 Pult
0.9 Pult
Pult
12000
Axial strain (με)
12000
Axial strain (με)
0.2 Pult 0.7 Pult
10000 8000 6000 4000 2000 0
10000 8000 6000 4000 2000 0
0
16
32
48
64
80
(x) distance from joint (mm) NA+0,1
96
0
16
32
48
64
80
(x) distance from joint (mm) NA+20,2
96
Figure
0.1 Pult 0.6 Pult
0.3 Pult 0.8 Pult
0.4 Pult 0.9 Pult
0.5 Pult Pult
10000 8000 6000 4000 2000
0.1 Pult 0.6 Pult
12000
Axial strain (με)
Axial strain (με)
12000
0.2 Pult 0.7 Pult
0.3 Pult 0.8 Pult
0.4 Pult 0.9 Pult
0.5 Pult Pult
10000 8000
6000 4000
2000 0
0 0
16
32
48
64
80
0
96
(x) distance from joint (mm) 0.1 Pult 0.6 Pult
12000
HA–40,1 0.2 Pult 0.3 Pult 0.7 Pult
0.8 Pult
0.4 Pult 0.9 Pult
16
32
48
64
80
96
(x) distance from joint (mm) 0.5 Pult Pult
0.1 Pult 0.6 Pult
HA–20,1 0.2 Pult 0.7 Pult
0.3 Pult 0.8 Pult
0.4 Pult 0.9 Pult
0.5 Pult Pult
12000
Axial strain (με)
Axial strain (με)
0.2 Pult 0.7 Pult
10000 8000 6000 4000
2000 0
10000 8000 6000 4000
2000 0
0
16
32
48
64
80
(x) distance from joint (mm) HA+0,2
96
0
16
32
48
64
80
(x) distance from joint (mm) HA+20,1
96
Figure
0.2 Pult 0.7 Pult
0.3 Pult 0.8 Pult
0.4 Pult 0.9 Pult
0.1 Pult 0.6 Pult
0.5 Pult Pult
40
Lap–shear stress (MPa)
Lap–shear stress (MPa)
0.1 Pult 0.6 Pult
30
20
10
0.3 Pult 0.8 Pult
0.4 Pult 0.9 Pult
0.5 Pult Pult
40
30
20
10
0
0 0
16
32
48
64
80
0
96
0.1 Pult 0.6 Pult
0.2 Pult 0.7 Pult
0.3 Pult 0.8 Pult
0.4 Pult 0.9 Pult
0.1 Pult 0.6 Pult
0.5 Pult Pult
Lap–shear stress (MPa)
40
30
20
10
0 0
16
32
48
64
80
(x) distance from joint (mm) NA+0,1
16
32
48
64
80
96
(x) distance from joint (mm) NA–20,1
(x) distance from joint (mm) NA–40,2
Lap–shear stress (MPa)
0.2 Pult 0.7 Pult
96
0.2 Pult 0.7 Pult
0.3 Pult 0.8 Pult
0.4 Pult 0.9 Pult
0.5 Pult Pult
40
30
20
10
0 0
16
32
48
64
80
(x) distance from joint (mm) NA+20,2
96
Figure
0.6 Pult
0.2 Pult
0.3 Pult
0.7 Pult
0.8 Pult
0.4 Pult 0.9 Pult
0.5 Pult Pult
40 30 20 10 0 0
16
32
48
64
80
96
Lap-shear stress (MPa)
Lap-shear stress (MPa)
0.1 Pult
HA–40,1 0.2 Pult 0.3 Pult
0.4 Pult
0.5 Pult
0.6 Pult
0.7 Pult
0.9 Pult
Pult
30 20 10 0 32
48
64
80
(x) distance from joint (mm) HA+0,2
0.4 Pult
0.5 Pult
0.8 Pult
0.9 Pult
Pult
30
20 10 0
0
16
32
48
64
80
96
96
HA–20,1 Lap-shear stress (MPa)
Lap-shear stress (MPa)
0.1 Pult
16
0.3 Pult
0.7 Pult
(x) distance from joint (mm)
40
0
0.2 Pult
0.6 Pult
40
(x) distance from joint (mm)
0.8 Pult
0.1 Pult
0.1 Pult
0.2 Pult
0.3 Pult
0.4 Pult
0.5 Pult
0.6 Pult
0.7 Pult
0.8 Pult
0.9 Pult
Pult
40 30
20 10 0 0
16
32
48
64
80
(x) distance from joint (mm)
HA+20,1
96
S1359-8368(14)00090-0 http://dx.doi.org/10.1016/j.compositesb.2014.02.013 JCOMB 2941
To appear in:
Composites: Part B
Received Date: Revised Date: Accepted Date:
13 October 2013 28 January 2014 14 February 2014
Please cite this article as: Al-Shawaf, A., Influence of fibres ’ stiffness on wet lay-up cfrp/steel joints ’ behaviour under subzero exposures, Composites: Part B (2014), doi: http://dx.doi.org/10.1016/j.compositesb.2014.02.013
This is a PDF file of an unedited manuscript that has been accepted for publication. As a service to our customers we are providing this early version of the manuscript. The manuscript will undergo copyediting, typesetting, and review of the resulting proof before it is published in its final form. Please note that during the production process errors may be discovered which could affect the content, and all legal disclaimers that apply to the journal pertain.
INFLUENCE OF FIBRES’ STIFFNESS ON WET LAY-UP CFRP/STEEL JOINTS’ BEHAVIOUR UNDER SUBZERO EXPOSURES Ahmed Al-Shawaf1 Department of Civil Engineering, Monash University, Victoria 3800, Australia [email protected] Department of Civil Engineering MONASH University 3800 Melbourne, Victoria, Australia T: +613 9879 7816 Fax: +613 9905 4944
Abstract The current paper characterizes and differentiates the bond behaviour of wet lay-up CFRP/steel double-strap joints fabricated with NM (normal modulus) and UHM (ultra high modulus) unidirectional CF (carbon fibres) plies. The influence of infrastructural subzero exposure on the bond attributes of these joints is also investigated. While environmental exposure is maintained at a particular subzero temperature, ranging from –40 to 20ºC, specimens representing these joints are tested in tension. Failure patterns, joints’ strength, and strain and LSS (lap-shear stress) distributions for the tested joints are extracted. Pertinent discussions and conclusions, related to the influence of CFRP moduli and subzero temperatures on the aforementioned parameters, are provided.
Keywords Cohesion failure; Load transfer; Normal modulus; Subzero exposures; Ultra high modulus; Wet lay-up
1
Corresponding author
1
1
Introduction Adhesively bonded CFRP/steel strap joints provide many advantages over the traditional steel-plate
bonding in retrofitting applications. These include remarkable strength and/or stiffness-to-weight ratio, faster field application due to its light weight, tailorable thermo-mechanical properties, relatively more durability as a material, higher chemical and corrosion resistance, and enhanced fatigue life [1]. In addition, it is a well-known fact that adhesively-bonded joints are strong in shear, but inevitably weak in peel. Therefore, double–strap joint configuration is chosen for the current work to ensure that the applied load is transferred mainly in shear and to minimize any direct or induced peel stresses which tend to detract from the joint’s capacity [2]. In CFRP retrofitting applications, CF reinforcement is commonly categorised into: (1) NM (Normal modulus) / high-strength, (2) HM (High modulus), and (3) UHM (Ultra high modulus). Ranges of CF elastic moduli for these categories generally fall into 230-240, 295-390, and 440-640 GPa, respectively [3]. However, the equivalent (i.e. overall) elastic moduli for the commonly-used preformed CFRP laminates varies from 150 to 450 GPa, respective to the lower and upper moduli limits of the aforementioned CF categories [4]. On the other hand, and for comparative CFRP layer’s thickness, wet/hand lay-up CFRP fabrication has considerably less range (e.g. 80 to 225 GPa [1]) due to practical limitations usually associated with this method. In this context, existing literature has a pronounced deficiency of pertinent wet lay-up strengthening publications compared to preformed methods [5-21]. Furthermore, there is even no universally accepted classification system, for the preformed methods, that researchers can refer to [22]. Therefore, it is proposed for the present work to ascribe wet lay-up CFRPs with moduli values below 100, 100-200, 200-400, and above 400 GPa as LM (Low Modulus), NM (Normal Modulus), HM (High Modulus), and UHM (Ultra High Modulus), respectively. It has been established that for any un-prestressed FRP strengthening material to be effective, its stiffness must be similar to the metallic structure being strengthened [3]. As such, NM CFRP is a cost-effective option for strengthening ‘carbon steel’ substrata as opposed to HM and UHM CFRP, since its strain and stress capacity can be mobilised after the substrate has yielded; and the strengthening material can be used considerably more effectively than with cast iron structures. On the contrary, cast iron is attributed with relatively low cracking strain, thus require FRP reinforcement with as high a modulus as possible to be effective in strengthening the section [3]. Consequently, HM and UHM CFRP would suit cast iron structures more efficiently. The majority of existing research on FRP strengthening for steel structures has mainly focused on preformed NM CFRP (i.e. CFRP modulus ≈ 100-200 GPa) [4, 22-26]. This research trend can be reasonably
2
justified by the higher number, worldwide, of infrastructural carbon steel structures requiring CFRP strengthening compared to the other less-encountered type of metallic structures made from cast iron. Accordingly, more research should be directed towards HM and UHM CFRP external strengthening applications. The other under-investigated issue, which affects the infrastructural behaviour of CFRP/steel joints, is the environmental thermal exposures under which CFRP strengthening systems operate. Part of these exposures is subzero environmental exposures commonly encountered in infrastructural CFRP strengthening applications. These exposures are not expected to alter, tangibly, the thermo-mechanical attributes of reinforcing CF plies. However, the HM CFRP strengthening material is most likely to reveal different load transfer and failure mechanisms relative to LM and NM CFRP, under any specific subzero thermal exposure. This in turn would be reflected on the overall joint’s attributes, such as ultimate strength, and interfacial strain and LSS distributions [27, 28]. In the present investigation, it is intended to investigate the effect of utilising NM and UHM CF plies in fabricating wet lay-up CFRP/steel plate double-strap specimens on their bond behaviour under subzero thermal exposures. In order to avoid under-impregnated CF plies usually encountered in the wet lay-up method, precisely with the UHM type, a special fabrication rig for producing wet lay-up CFRP/steel specimens is adopted. The same steel adherend (i.e. carbon steel) shall be utilized for both CF types as the aim is to conduct bond behavioural comparison, with as minimum variables as possible. These specimens shall be tested in direct tension, under short-term sustained subzero environmental exposures ranging from –40 to 20°C. Assessments of failed specimens, at any subzero temperature, shall be based on their concluded failure patterns, capacities, longitudinal strain and LSS distribution within the joint’s interfacial adhesive layer.
2
Testing program As mentioned earlier, double-strap joint with thin adherend (i.e. 4.85 mm steel plate) is adopted, as
this provides optimum configuration to ensure that the applied load is transferred mainly in mode II (i.e. shear), and to minimize any direct or induced peel stresses. Three CF plies were adopted as reinforcement for the wet lay-up CFRP laminate on each face of the double-strap joint. This number was concluded to be the optimum, in terms of load transfer efficiency, based on a previous investigation which adopted similar CF plies with wet layup method on steel plates [29]. The variables of the current specimens are: (1) elastic modulus of the CF plies
3
(i.e. NM and UHM), and (2) subzero temperatures ranging from –40 to 20°C simulating thermal exposures encountered in outdoor CFRP retrofitting applications.
2.1
Wet lay-up joint’s constituents and processing The two unidirectional NM and UHM CF plies adopted for the current investigation are MBrace®CF
130 and 530, respectively. Both CF types were provided by Degussa Construction Chemicals Australia Pty Ltd. Their nominal thickness is 0.176 and 0.19 mm, respectively; and thermo-mechanical properties are quoted Table 1 according to the manufacturer’s technical data sheets [30, 31] and some literature standard-products data, as cited in the same table. Araldite®420 A/B is a commercial epoxide resin utilized as the matrix/adhesive for the present wet lay-up CFRP/steel double-strap joints, for both NM and UHM CF plies’ specimens. It was provided by Huntsman Advanced Materials. Its experimentally-extracted thermo-mechanical properties at current subzero environmental exposures are presented in Table 2.
Experimental procedures for characterising resin’s
parameters are detailed in Al-Shawaf [1]. Only the coefficient of thermal expansion was obtained from relevant adhesive’s literature (e.g. manufacturer data sheet). For the steel adherend, mild steel plates with 380 MPa yield strength, 227 GPa Young’s modulus, and nominal thickness of 4.85 mm are used in the current program. The steel plate was purchased from OneSteel Australia Limited. The wet lay-up CFRP/steel plate double-strap specimens were fabricated with a specially-designed steel rig shown in Figure 1. For detailed CFRP samples’ fabrication and rig’s design justification and advantages, readers are referred to Al-Shawaf and Zhao [33]. Each CFRP/steel sample was left for 7 days to cure under ambient exposure prior to water-jet cutting into 8 identical double-strap specimens, according to the predetermined dimensions.
2.2
Specimen setup, thermal conditioning, and testing ERSGs (electrical resistance strain gauges), compatible with plastic surfaces and with wide-range
limits of operational temperature from –60 to +250°C, were instrumented on the CFRP outer surface, as shown in Figure 2, to capture axial strain readings along the bondlength during load application. The application of the ERSGs and specimens’ setup, thermal conditioning, and testing procedures are detailed in Al-Shawaf and Zhao [33].
4
2.3
Current CFRP/steel joints’ configurations Figure 3 (a and b) indicates the two configurations of the FRP joints adopted in this study. The details
of this figure are self-explanatory. As can be implied, the modulus for the wet lay-up CFRP strengthening material for the whole joint is twice that on each face of the double-strap configuration. Accordingly, the current NM CFRP/steel joint (i.e. Figure 3a) is anticipated to be more effective compared to the UHM CFRP/steel joint of Figure 3b, because its overall stiffness (i.e. ≈ 160 MPa) is considerably closer to that of the steel plate compared to the UHM CFRP/steel joint. Table 3 below and the following paragraph, facilitates the interpretation of the results, discussions, and analyses presented later. Specimens ID labels reflect the material configuration and exposure temperature for each of the experimental composite joints. The current subgroups are NA and HA. The first character denotes the modulus grade of the adopted CF plies, viz. (N) for NM CF (i.e. E = 240 GPa) and (H) for UHM CF (i.e. E = 640 GPa). The following (A) character acronyms the Araldite adhesive/matrix. Then, the environmental exposure temperature follows. Lastly, the replicated number of any particular specimen is indicated. For example, HA20,2 denotes the second specimen fabricated with UHM CF and conditioned at -20°C.
3
Analyses of results Modes of failure, joints’ capacities, and interfacial longitudinal strain and LSS distributions, for all
tested CFRP/steel double-strap joints, are discussed in this section. Informative analyses on these parameters are properly addressed and justified. Comparative assessments of the effect of subzero temperatures on both types of specimens (i.e. NM and UHM CFRP/steel joints) are conducted, based on the abovementioned parameters.
3.1
Modes of failure Figure 4 is deemed useful in interpreting current failure modes. It depicts the likely modes of failure
encountered in FRP/ steel plate double-strap joints. As a prelude, any monotonically-loaded CFRP/steel joint upon failure displays one of two apparent failure modes’ types. These can be described as ‘single-mode’ and ‘mixed-mode’. In the former category, as the name implies, failure pattern is manifested wholly by any single mode of Figure 4. Practically, this case is rare; therefore the ‘mixed-mode’ is frequently prevailing. Nonetheless, few exceptions apply, such as bond failures at extremely elevated temperatures’ exposures close to the interfacial adhesive’s Tg (glass transition temperature),
5
or if the interfacial adhesive/matrix is incompatible with steel adherends [1]. In mixed-mode failures, and with most commercially steel-compatible adhesives/matrices, actual bond capacity for well-configured CFRP/steel joints is closely related to the failure mode responsible for failure initiation (i.e. triggering failure). FEA (finite element analysis) method has proved to provide effective support in acquiring reasonable pre-experience required to determine the most likely triggering mode of failure [28]. Figure 5 assists the reader, at later discussions, in referring to the 4 potential sides (i.e. bondlengths) of the current joints’ failure locations.
3.1.1
NA-subgroup All specimens of this subgroup, at current subzero thermal exposures, exhibited mixed-mode failure
type of varying intensity, as indicated in Figure 6. The predominant failure mode was interlaminar failure in the CFRP laminate (i.e. Figure 4d). Primarily, this implies that the current fabrication rig was successful in providing quality adhesion and preventing premature interfacial debonding failure (i.e. Figure 4a). Moreover, ‘localized interfacial debonding’ (i.e. type (a) of Figure 4) were detected in the proximity of either/both ends of the instrumented bondlength. This pattern was found in all NA-subgroup specimens. It was found in Al-Shawaf [1], upon performing FEA of these joints, that their triggering failure pattern is ‘cohesion failure’ similar to Figure 4c and located close to either end, as well. As such, it is believed that current failures were triggered within the interfacial adhesive layer by a tensile cohesion failure close to either end of the bondlength creating scattered deep hackles, which were then propagated and eventually linked up at the adhesive/steel interface before the CF/adhesive interface due the continuum nature of the steel surface, as opposed to the filamentary nature of the carbon fibres. Hence, those localized interfacial debonding patches were observed, experimentally. For more details on this failure mechanism, readers are referred to Hart-Smith [2] / Al-Shawaf and Zhao [33]. The latter failure pattern for the NA-subgroup specimens has not been influenced by the current range of subzero exposure temperatures (i.e. –40, –20, and 0°C). The only observed difference was the increased brittleness and violent nature of the failure process as the conditioning exposure temperature of the specimens decreased.
3.1.2
HA-subgroup Tested specimens belonging to this subgroup are portrayed in Figure 7.
6
The apparent predominant failure mode for the HA-subgroup specimens is CFRP rupture at the joint (refer to Figure 4e). Also, very small areas (about 2% of the overall bond area for the four bondlengths) of ‘localized interfacial debonding’ (i.e. type (a) of Figure 4), close to the 2 mm gap were visually detected, especially for the 20°C-exposure specimens. This failure pattern can be attributed as ‘fibre-controlled’, as opposed to the ‘adhesive-controlled’ failure pattern found in the NA-subgroup specimens. The dissimilarity of the current failure pattern from the NA-subgroup is related to the difference in joints’ load transfer mechanism, and thus interfacial adhesive stress distribution. This different load transfer mechanism is basically a direct consequence of lower tensile strength and ultimate elongation values (i.e. 2070 MPa and 0.4%, respectively) of the UHM CF plies compared with their respective NM CF plies’ values of the NA-subgroup, viz. 3800 MPa and 1.55%, respectively. It was found, via FEA for the HA-subgroup joints that although 1st CF ply’s rupture at the joint initiated failure, its propagation would not have been facilitated had the strain energy level of the neighbouring interfacial adhesive layer been relatively low. Moreover, FEA in the vicinity of the mid-gap showed that HA–40 specimens had simultaneously attained ultimate capacities for the 1st CF ply and the interfacial adhesive/matrix material [1]. Accordingly, it can be concluded that the common instigator of the HA-subgroup joints’ triggering failure is related, either directly or indirectly, to the interfacial adhesive layer, viz. its stress status. As with the NA-subgroup, current subzero thermal exposures had almost negligible effect on the predominant failure mode trend.
3.2
Ultimate joints’ capacity Table 4 presents the ultimate loads for current specimens. In order to be consistent in concluding the
short-term effect of environmental exposures on joint capacity, the triggering failure pattern should be identical for all specimens involved in any comparison representing the whole temperature spectrum. From Subsection 03.1, this prerequisite is found applicable for the present work.
3.2.1
NA-subgroup Based on the preceding paragraph, all specimens of the NA-subgroup developed identical failure
patterns, both triggering and predominant. As such, noting their corresponding average ultimate loads at current subzero and ambient exposures in Table 4, no special trend was inferred from increasing the short-term exposure temperature from –40°C to 20°C. However, by comparing the relevant joints’ capacities with those
7
extracted from corresponding simulated FE models [28], close validation was concluded for the FE models of the NA-subgroup specimens with the higher experimental capacities amongst the two replicates of each exposure (i.e. NA–40,1, NA–20,1, NA+0,1, and NA+20,2). Consequently, and taking into consideration the expected experimental results’ scatter between replicated specimens, the ones with higher bond capacities can be assumed as representative of the accurate behaviour at each environmental exposure. The ultimate capacities of those specimens suggest that there is a reduction of about 4.5% for each 20°C temperature increase within the current subzero temperatures’ range , which is generally in line with few previous investigations [5, 34].
3.2.2
HA-subgroup The generic ultimate load capacity trend for specimens of this subgroup is characterized by lower
values than the NA-subgroup. Apparently, the direct reason for this observation is the considerably stiffer and lower tensile strength of the UHM CF plies of the HA-subgroup specimens as compared to the NA-subgroup counterparts. The latter trend is particularly evident at the subzero temperatures. From Table 4, it can be concluded that there is no tangible difference between the HA-subgroup average joints’ capacities at –40 and –20°C exposures; however they increased by 17% and 5% at 0 and +20°C exposures, respectively. It is believed that the abovementioned variations are related to the different contributions of the adhesive/ matrix in the load transfer mechanism. More precisely, the substantial 35% increment in the adhesive’s ‘elongation at failure’ as the conditioning temperature increased from –20 to 0°C is a reasonable justification of the aforementioned 17% joint’s capacity increment. This 35% extra elongation capacity simply enhanced the interfacial adhesive’s ability of accommodating more shear lag deformations before failure. Consequently, the average joint’s LSS, and thus ultimate capacity, have increased. The same scenario is proportionately applicable as the thermal exposure increased from 0 to 20°C. This behaviour clearly highlights the role played by the rheological properties of the adhesive/matrix in determining the individual load transfer mechanism, at each subzero environmental exposure, and its key influence in shaping the joint’s failure pattern and ultimate capacity.
3.3
Axial strain distributions Strain readings along the instrumented bond length were captured via standard ERSGs (number 1 to 7
as indicated in Figure 2). The total and gauge lengths for the current ERSGs are 13 and 6 mm, respectively. It has been demonstrated that, with standard ERSGs, some experimental / practical strain-measuring constraints have to be considered in order to meaningfully assess the accuracy of strain readings, and thus LSS. In this
8
study, these strain-measuring constraints exist at bondlength ends (i.e. locations of SGs 1 and 7). Accurate and representative strain readings at these locations are paramount due to their close relevance to joint’s triggering failure proximity. Close to x = 0, strain/stress exhibits sharp peak away from directly above the 2-mm gap (Figure 2) due to some limited bending of the CFRP reinforcement at this point. This bending, which is undetectable by SG 1 due to its 6-mm gauge length (i.e. it is actually overlapping the 2-mm gap), is believed to cause a compressive stress/strain at the surface (i.e. outer side) of the CFRP layer within this narrow proximity [28]. At the other end (i.e. exactly at x = 100 mm), it is practically impossible to obtain strain readings at the CFRP’s edge with traditional strain gauging techniques. For these reasons, and more broadly, the complexity of the stress distributions in bonded joints; their theoretical analyses is very difficult to verify experimentally, which renders FEA methods as the optimum recommended option for accurate behavioural predictions within the interfacial adhesive layer [35]. Apart from bondlength ends, accurate experimental strain distributions were obtained for both NA and HA-subgroups, from x = 16 to 96 mm. These where validated via FEM predictions of the current experimental joints [1]. It has been assumed that the inherent existence of shear lag deformations within CFRP laminate’s thickness is ignored due to the relatively stiff response of most structural-grade adhesives at subzero to ambient thermal exposures [1]. In addition, it is assumed that bulk longitudinal and shear deformations in the adherend (i.e. steel plate) can also be neglected due to their small values relative to those of the wet lay-up CFRP laminate and adhesive layer.
3.3.1
NA-subgroup Figure 8 depicts the axial strain readings, for one replicate from each thermal exposure of the NA-
subgroup specimens, plotted with distance from the joint, for 10% successive increments of ultimate failure load. The general pattern in the strain distributions for subzero and ambient exposures is, as expected, ultimate strain values at the joint (i.e. x = 0) followed by gradual reductions towards the other end of the bondlength at all load levels (i.e. from 10% Pult. to Pult.); with the rate of strain reduction decreasing sharply from x = 0 to x = 32 mm, after which almost uniform distributions are displayed up to x = 80 mm, then approaches zero at x = 96 mm (i.e. location of SG7). This trend is consistent with the traditional stiffness-imbalanced and thermally-mismatched doublelap joint model [36]. The latter model establishes that, for a double lap-shear joint, whenever the net stiffness at one end of a joint differs from that at the other end, the adhesive shear strain distribution is rendered non-
9
symmetric. The second cause of strain non-symmetry is the thermal mismatch between adherends, according to the same model. In case of differences in coefficients of thermal expansion between dissimilar adherends, more severe strains are associated with the end where the lower coefficient of thermal expansion’s adherend extends. From Al-Shawaf [1], the coefficients of thermal expansion for the NA-subgroup CFRP laminate are 2.408, 2.042, 1.620, and 1.521x10–6/°C at –40, –20, 0, and 20°C thermal exposures, respectively; and for steel is 11x10-6/°C. Therefore, thermal mismatch effect has actually added up shear strains close to x = 0 end of the joint. Comparison of the 4 specimens in Figure 8 reflects the anticipated effect of increasing thermal exposure from –40°C to 20°C. This effect is manifested in slightly increasing both the uniformity of strain distributions along the bondlength and the ultimate strain values at the joint. The latter trends are direct consequences of decreasing embrittlement (i.e. increasing ductility) with temperature for joints with ductile adhesives, such as the current one.
3.3.2
HA-subgroup The axial strain distributions for selected HA-subgroup specimens, representing all currently-
experimented thermal exposures, are presented in Figure 9. In terms of the general trend, it resembles the previously-discussed one in that the highest strain is always at x = 0, followed by gradual declination in strain values towards the other end of the instrumented bondlength (i.e. x = 100 mm). Two outstanding attributes can be easily concluded from the comparison of Figures 9 and 8. The first one is the negligible difference amongst the ultimate strain values for the four specimens of the HA-subgroup compared to the respective tangible strain increments with conditioning temperatures for the NA-subgroup. This aspect supports the ‘fibre-controlled’ load transfer mechanism of the former subgroup as opposed to the ‘adhesive-controlled’/ ‘adhesive dependant’ for the latter. The second is the approximately 78% drop in the ultimate strain values at the joint (i.e. x = 0) exhibited by the HA-subgroup specimens relative to the NAsubgroup ones which clearly reflects the considerable stiffness’s difference between both CFRPs.
3.4
Lap-shear stress distributions The average LSS between two adjacent ERSGs was calculated by dividing the force difference by the
bond area between both gauges (per unit width of the bond), according to the following equation:
τ = E frp . (ε f i +1 − ε f i ). t frp / Δl
10
(1)
where Efrp and tfrp are the CFRP elastic modulus and thickness, respectively; εfi+1 and εfi are the CFRP axial strains (i.e. strain gauge readings) at two consecutive locations Δℓ distance apart (i.e. 16 mm as indicated in Figure 2). Stresses calculated via Equation (1) actually represent the average shear stresses at 16 mm intervals along bondlength, thus are smaller than actual values in the composite specimen. A dominant trend of very good agreement was detected between experimental LSS distributions and their predicted FE counterparts, from x = 24 to 88 mm along instrumented bondlength, for both subgroups and at all load levels [1]. Locations closer to both bondlength ends revealed some divergence from the latter trend due to practical reasons explained earlier (i.e. in the prelude to Section 3.33.3).
3.4.1
NA-subgroup LSS distributions along instrumented bondlength for one replicate of the NA-subgroup specimens are
presented in Figure 10. The general trend was successfully verified against existing analytical model/s in AlShawaf [1]. The established trend of LSS distribution in a balanced double-lap joint (i.e. Eiti = 2Eoto) is the nonuniformity of the load transfer throughout the adhesive bondlength, which peaks at each end of the overlap and is symmetrical about a perpendicular line passing through x = ℓ/2 of the bondlength [2]. In the current investigation, due to adherend stiffness imbalance and the thermal mismatch of all double-strap joint specimens, the aforementioned symmetry is not the case. Instead, the prevailing trend for the NA-subgroup specimens under all environmental exposures is the non-symmetry in LSS distribution, with peak values occurring close to x = 0 end. From Figure 10, the embrittlement effect at subzero temperatures is clearly dominating the LSS behaviour of specimens NA–40,2, NA–20,1, and NA+0,1 as compared to specimen NA+20,2. It can also be detected that the –40, 0, and 20°C exposure specimens have followed similar trend close to x = 0 end of the bondlength, and for almost all load levels where the adhesive behaviour is linear. The only exception to this trend occurred with both specimens NA–20,1 and 2. This distinct LSS trend, at –20°C exposure, can be traced back from the corresponding axial strain distribution in Figure 8. Apparently, this could be misinterpreted as adhesive non-linearity, however close examination of specimen’s NA–20,1 LSS distribution (refer to Figure 10) reveals comparable stress values at both x = 8 and 24 mm locations, even at low-load levels where adhesive usually behaves linearly. The likely explanation to this phenomenon is believed to be due to different thermal residual stress’s response for the CFRP laminate at each conditioning temperature. These residual thermal stresses add up to the mechanical stresses to shape up the final bond LSS distribution.
11
3.4.2
HA-subgroup Figure 11 portrays the interfacial LSS distributions along the instrumented bondlengths of four
representative specimens from HA-subgroup. It is believed that LSS distributions are of secondary significance in CFRP joints where load transfer and failure mechanisms are driven by the reinforcing fibres (i.e. fibrecontrolled), as the case for HA-subgroup specimens. As such, the following discussion is merely to stress the behavioural difference of LSS distributions between both NA and HA-subgroups. The same characteristic LSS distribution trend of peak values close to bondlength ends is applicable for the specimens of this subgroup. However, the key variation reflected on the HA-subgroup’s specimens relative to the NA subgroup’s is the conversion of the LSS peaks from ‘x = 0’ end, for all specimens of the latter subgroup, to ‘x = 100 mm’ end for the corresponding specimens of the former subgroup. This variation was apparently by virtue of the difference occurring to the combined effect of adherends’ stiffness imbalance and thermal mismatch. The other useful implication extracted from Figure 11 is related to the HA-subgroup specimens’ strengths compared to the NA-subgroup ones. From both Figure 10 and Figure 11, the average LSS of specimens belonging to the former is apparently higher than that for correspondent specimens of the latter. This finding is useful in predicting/justifying the higher joints values, presented earlier, for the NA-subgroup specimens compared to their counterparts from the HA-subgroup. In terms of the effect of subzero exposures, no clear trend can be implied on the LSS distributions of the HA-subgroup specimens. The ‘fibre-controlled’ behaviour of the current specimens is anticipated to, reasonably, justify this trend.
4
Conclusions The prime conclusion inferred from the current study is the substantial role of the adhesive / matrix
for wet lay-up CFRP/steel joints reinforced with NM CF plies in the load transfer process, as opposed to the predominance of the CF plies’ role for similar joints, however, reinforced with UHM CF plies. More specific parameter-related conclusions for the wet lay-up fabrication method are drawn in the following subsections:
4.1
Failure modes
• Current subzero environmental exposures have almost negligible effect on the individual predominant failure mode for double-strap NM CFRP/steel joints and UHM CFRP/steel joints.
12
• Double-strap NM CFRP/steel joints manifest ‘adhesive-controlled’ characteristic failure mode; while UHM CFRP/steel joints reveal clear ‘fibre-controlled’ mode. • For ductile interfacial adhesives, the likely triggering failure pattern for double-strap NM CFRP/steel joints is cohesion failure in the adhesive layer close to either ends of the bondlength, and CFRP rupture for UHM CFRP/steel joints.
4.2
Bond strengths
• Double-strap CFRP/steel joints manifesting ‘Adhesive-controlled’ behaviour usually attain higher capacities than identical joints disclosing ‘fibre-controlled’ behaviour. • Capacities for UHM CFRP/steel joints are more susceptible to variation in subzero thermal exposures within current range than NM CFRP/steel joints fabricated with identical ductile interfacial adhesive.
4.3
Axial strain distributions
• For CFRP/steel joints fabricated with identical ductile adhesive, ultimate strain values are inversely proportional with CF plies’ stiffness. • Axial strain distributions for NM CFRP/steel joints are more vulnerable to thermal variation within current subzero exposures than UHM CFRP/steel joints. • Standard ERSGs are incapable of capturing accurate strain values at bondlength ends where load transfer is highest. Instead, innovative strain-measuring techniques should be researched, such as FOS (fibre optic sensors), advanced non-contact optical and scanning electron microscopy (SEM) methods.
4.4
LSS distributions
• Measuring interfacial LSS distributions experimentally is more useful in predicting failure of CFRP/steel joints attributed with ‘adhesive-controlled’ behaviour; as opposed to ‘fibre-controlled’ behaviour’s joints. • The generic LSS distribution trend conforms to the traditional model of linear distribution with its characteristic peak values close to both ends of the bondlength at loads below the triggering failure one. • Ultimate LSS’s location is function of the CFRP layer’s stiffness. It is close to x = 0 mm in joints with low CFRP stiffness, as opposed to x = ℓ for joints with high CFRP stiffness. • As with the strain distribution trends, LSS distributions for NM CFRP/steel joints are more susceptible to thermal variations, within current subzero exposures, than UHM CFRP/steel joints.
13
5
Acknowledgements This work is sponsored by an Australian Research Council (ARC) Discovery Grant. The kind
assistance of Jim Mitchell and Silvio Mattievich in setting up and operating the environmental chamber and the Instron testing machine at the Department of Materials Engineering / Monash University is acknowledged by the author. He also wishes to thank Kevin Nievaart for his contribution in facilitating the progress of the experimental part at both the Civil and Instron laboratories. Finally, the great help of all staff at the Civil Laboratory / Monash University is appreciated in executing the experimental program.
References 1. Al−Shawaf A. Characterization of bonding behaviour between wet lay-up carbon fibre reinforced polymer and steel plates in double-strap joints under extreme environmental temperatures. Doctorate PhD thesis, Monash University, Melbourne, Australia, 2010. 2. Hart−Smith LJ. Bolted and bonded Joints. In: Miracle DB, Donaldson SL, editors. Composites. Ohio, USA: ASM International, 2001. p.271-89. 3. Cadei JMC, Stratford TJ, Hollaway LC, Ducket WG. Strengthening metallic structures using externally bonded fibre-reinforced polymers C595. London: CIRIA, 2004. 4. Zhao XL, Zhang L. State of the art review on FRP strengthened steel structures. Eng Struct 2007;29(8): 180823. 5. El−Hacha R, Green MF, Wight G. Strengthening of concrete members with prestressed FRP sheets: Behaviour at room and low temperature. London, Ont., Canada: Canadian Society for Civil Engineering, Montreal, H3H 2R9, Canada, 2000:187. 6. di Tommaso A, Neubauer U, Pantuso A, Rostasy FS. Behavior of adhesively bonded concrete-CFRP joints at low and high temperatures. Mech Compos Mater 2001;37(4): 327-38. 7. Bisby LA, Green MF. Resistance to freezing and thawing of fiber-reinforced polymer-concrete bond. ACI Struct J 2002;99(2): 215-23. 8. Colombi P, Fanesi E, Fava G, Poggi C. Durability of steel elements strengthened by FRP plates subject to mechanical and environmental loads. CCC2005 : Third International Conference Composites in Construction. Lyon, France, 2005. 9. Green MF. FRP repair of concrete structures: performance in cold regions. International journal of materials and product technology 2007;28(1-2): 160.
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10. Karbhari VM, Rivera J, Dutta PK. Effect of short-term freeze-thaw cycling on composite confined concrete. J Compos Constr 2000;4(4): 191-97. 11. Abanilla MA, Karbhari VM, Li Y. Interlaminar and intralaminar durability characterization of wet layup carbon/epoxy used in external strengthening. Compos Part B-Eng 2006;37(7-8): 650-61. 12. Agarwal A, Foster SJ, Hamed E, Ng TS. Influence of freeze–thaw cycling on the bond strength of steel–FRP lap joints. Compos Part B-Eng 2014;60(0): 178-85. 13. Al−Emrani M, Kliger R. Analysis of interfacial shear stresses in beams strengthened with bonded prestressed laminates. Compos Part B-Eng 2006;37(4-5): 265-72. 14. Almusallam TH, Al-Salloum YA. Ultimate strength prediction for RC beams externally strengthened by composite materials. Compos Part B-Eng 2001;32(7): 609-19. 15. Camata G, Spacone E, Zarnic R. Experimental and nonlinear finite element studies of RC beams strengthened with FRP plates. Compos Part B-Eng 2007;38(2): 277-88. 16. Colombi P, Poggi C. An experimental, analytical and numerical study of the static behavior of steel beams reinforced by pultruded CFRP strips. Compos Part B-Eng 2006;37(1): 64-73. 17. Czaderski C, Rabinovitch O. Structural behavior and inter-layer displacements in CFRP plated steel beams – Optical measurements, analysis, and comparative verification. Compos Part B-Eng 2010;41(4): 276-86. 18. Keller T, Vallée T. Adhesively bonded lap joints from pultruded GFRP profiles. Part I: stress–strain analysis and failure modes. Compos Part B-Eng 2005;36(4): 331-40. 19. Kim YJ, Harries KA. Behavior of tee-section bracing members retrofitted with CFRP strips subjected to axial compression. Compos Part B-Eng 2011;42(4): 789-800. 20. Parvin A, Granata P. Investigation on the effects of fiber composites at concrete joints. Compos Part B-Eng 2000;31(6–7): 499-509. 21. Yu T, Fernando D, Teng JG, Zhao XL. Experimental study on CFRP-to-steel bonded interfaces. Compos Part B-Eng 2012;43(5): 2279-89. 22. Peiris NA. Steel beams strengthened with ultra high modulus CFRP laminates. Doctorate PhD thesis, University of Kentucky, Lexington, 2011. 23. Zhao XL. Recent developments in FRP strengthening of metallic structures. 1st Middle East conference on smart monitioring, assessment and rehabilitation of civil structures. Dubai, UAE, 2011. 24. Yu Y, Chiew SP, Lee CK. Bond failure of steel beams strengthened with FRP laminates – Part 2: Verification. Compos Part B-Eng 2011;42(5): 1122-34.
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25. Sen R, Liby L, Mullins G. Strengthening steel bridge sections using CFRP laminates. Compos Part B-Eng 2001;32(4): 309-22. 26. Linghoff D, Al-Emrani M, Kliger R. Performance of steel beams strengthened with CFRP laminate – Part 1: Laboratory tests. Compos Part B-Eng 2010;41(7): 509-15. 27. Adams RD. Adhesive bonding : science, technology and applications. Cambridge: Woodhead Publishing Ltd, 2005:xvi, 543. 28. Al−Shawaf A. Modelling wet lay-up CFRP–steel bond failures at extreme temperatures using stress-based approach. Int J Adhes Adhes 2011;31(6): 416-28. 29. Fawzia S, Al−Mahaidi R, Zhao XL, Rizkalla SH. Strengthening of circular hollow steel tubular sections using high modulus CFRP sheets. Constr Build Mater 2007;21(4): 839-45. 30. BASF. MBrace® CF130 technical data sheet www.basf-cc-la.com, 2006. 31. BASF. MBrace® Fibre technical data sheet www.basf-cc.com.au, 2006. 32. BP−Amoco. Thornel carbon fibres-http://www.cytec.com/engineered-materials/downloads/ThornelTP.pdf. Available from URL: http://www.cytec.com/engineered-materials/downloads/ThornelTP.pdf. 33. Al-Shawaf A, Zhao XL. Adhesive rheology impact on wet lay-up CFRP/steel joints’ behaviour under infrastructural subzero exposures. Compos Part B-Eng 2013;47(0): 207-19. 34. Karbhari VM, Shulley SB. Use of composites for rehabilitation of steel structures - determination of bond durability. J Mater Civil Eng 1995;7(4): 239-45. 35. Harris JA, Adams RD. Strength prediction of bonded single lap joints by non-linear finite element methods. Int J Adhes Adhes 1984;4(2): 65-78. 36. Hart−Smith LJ. Adhesive-bonded double-lap joints. Hampton, Virginia, USA: NASA Langley Research Centre, 1973:106.
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6
Figure Captions
Figure 1 CFRP/steel double-strap joint fabrication rig, (a) schematic section [not-to-scale], and (b) photo of assembled rig Figure 2 Geometry and instrumentation of experimental program’s specimens (Not-to-scale) Figure 3 The two currently-adopted wet lay-up CFRP/steel double-strap joints’ configurations (not-to-scale) Figure 4 Possible modes of failure for double-strap configuration of FRP/steel adherend joint Figure 5 Schematic key diagram for designating the four potential failure sides for the current double-strap joint configuration Figure 6 Failure patterns for the NA-subgroup specimens Figure 7 Failure patterns for the HA-subgroup specimens Figure 8 Axial strain distributions along instrumented bondlength for NA-subgroup specimens Figure 9 Axial strain distributions along instrumented bondlength for HA-subgroup specimens Figure 10 LSS distributions along instrumented bondlength for NA-subgroup specimens Figure 11 LSS distributions along instrumented bondlength for HA-subgroup specimens
17
7
Tables
Table 1 Nominal properties for MBrace®CF 130 and 530 carbon fibre plies CF category MBrace®CF 130 MBrace®CF 530 * [19]
Long. Specific thermal heat conductivity (j/Kg.°C) (j/s/mm/°C)
Modulus of elasticity (GPa)
Long. strength (MPa)
Ultimate elongation (%)
Shear strength (MPa)
Long. Poisson’s ratio
Long. CTE (x 10-6/°C)
240
3800
1.55
117.2 *
0.27 *
-0.38
9.38
711 *
640
2070 *
0.4
50 *
0.245 *
-0.83
68.7
711 *
18
Table 2 Typical thermo-mechanical properties for Araldite®420 A/B epoxide resin Thermal Tensile exposure modulus (MPa) (°C) –40 –20 0 20
2974 2575 2118 2012
Poisson’s ratio
Tensile strength (MPa)
Elongation at failure (%)
0.364 0.346 0.333 0.366
33.26 43.90 40.17 32.60
1.24 1.99 2.68 2.85
Coefficient of Glass Thermal thermal conductivity transition expansion [W/m.°C] temp. (Tg) [°C] [x10-6/°C] 0.1431
19
41.66
110
Specific heat (j/g.°C) 0.8985 0.9679 1.0207 1.1328
Table 3 Current CFRP/steel double-strap specimens IDs Adopted CF category ®
MBrace CF 130 MBrace®CF 530
–40 NA–40,1 NA–40,2 HA–40,1 HA–40,2
Environmental exposure (°C) –20 0 NA–20,1 NA+0,1 NA–20,2 NA+0,2 HA–40,1 HA+0,1 HA–40,2 HA+0,1
20
20 NA+20,1 NA+20,2 HA+20,1 HA+20,2
Table 4 Ultimate loads of the current CFRP/steel double-strap specimens NA-subgroup Joint Specimen capacity (kN) NA–40,1 94.8 NA–40,2 75.6 NA–20,1 89.3 NA–20,2 77.3 NA+0,1 88.9 NA+0,2 76.5 NA+20,1 82.5 NA+20,2 88.7
Average capacity (kN) 85.2 83.3 82.7 85.6
HA-subgroup Joint Average Specimen capacity capacity (kN) (kN) HA–40,1 58.9 58.5 HA–40,2 58.0 HA–20,1 69.1 58.1 HA–20,2 47.1 HA+0,1 70.5 68.1 HA+0,2 65.6 HA+20,1 78.8 71.7 HA+20,2 64.5
21
8
Figures’ legends
•
Figure 8 Legend (identical for the 4 specimens)
•
Figure 9 Legend (identical for the 4 specimens)
•
Figure 10 Legend (identical for 4 specimens)
•
Figure 11 Legend (identical for 4 specimens)
22
Figure
Two screw-driven press plates
Press supporting frame
Steel angle
Wet Lay-up CFRP laminates (4.85 mm) Two sample steel plates
Steel angle
(2mm) Shims plates
(a)
(b)
Figure
4.85 mm
2 Wet lay-up CFRP layers
S.G.9
2 mm 2 mm
S.G.8
2-mm gap
steel plate
steel plate
S.G.1 S.G.2 S.G.3 S.G.4 S.G.5 S.G.6 S.G.7
Side view Total specimen length = 302 mm 6 @ 16.0 mm
Width = 50 mm S.G.1 S.G.2 S.G.3 S.G.4 S.G.5 S.G.6 S.G.7
steel plate
S.G.8
Bottom face view 100 mm
100 mm S.G.9
wet lay-up CFRP length = 202mm
Top face view
steel plate
Figure
LM CFRP laminate Ecfrp ≈ 80 GPa
4.85 mm
2 mm
NM CF-3 plies Eply = 240 GPa
2 steel plates
Identical bottom wet LM CFRP laminate (a) NM CFRP/steel double-strap joint (i.e. total CFRP modulus ≈ 160 GPa)
HM CFRP laminate Ecfrp ≈ 225 GPa 2 mm
UHM CF-3 plies Eply = 640 GPa
4.85 mm
Identical bottom wet HM CFRP laminate (b) UHM CFRP/steel double-strap joint (i.e. total CFRP modulus ≈ 450 GPa)
Figure
FRP laminate
Adhesive
(a) Substrate-adhesive debonding
(b) Adhesive-FRP debonding
(c) Cohesion failure
(d) Interlaminar failure in the FRP laminate
(e) FRP rupture at the joint
Figure
Side C
Side A
Side D
Side B
Figure
Failure sides: A + D Failure pattern: localized type (a) + type (d)
a
Failure sides: A + B Failure pattern: localized type (a) + type (d) a
Failure sides: A + D Failure pattern: localized type (a) + type (d) a
Failure sides: A + B Failure pattern: localized type (a) + type (d) a
Failure sides: A + D Failure pattern: localized type (a) + type (d) a
Failure sides: A + B Failure pattern: localized type (a) + type (d) a
Failure sides: A + B Failure pattern: localized type (a) + type (d) a
Failure sides: A + B Failure pattern: localized type (a) + type (d) a
a
Refer to Figure 5 for interpretation of failure mode types.
Figure
Failure sides: CFRP rupture (B/D) + C Failure pattern: type (e) + type (d) a
Failure sides: Total CFRP rupture Failure pattern: type (e) a
Failure sides: Total CFRP rupture Failure pattern: type (e) a
Failure sides: Total CFRP rupture Failure pattern: type (e) a
Failure sides: Total CFRP rupture Failure pattern: type (e) a
Failure sides: Total CFRP rupture Failure pattern: type (e) a
Failure sides: Total CFRP rupture Failure pattern: type (e) a
Failure sides: Total CFRP rupture Failure pattern: type (e) a
a
Refer to Figure 5 for interpretation of failure mode types.
Figure
0.2 Pult 0.7 Pult
0.3 Pult 0.8 Pult
0.4 Pult 0.9 Pult
0.1 Pult 0.6 Pult
0.5 Pult Pult
12000
12000
10000
10000
Axial strain (με)
Axial strain (με)
0.1 Pult 0.6 Pult
8000 6000 4000 2000
0.3 Pult 0.8 Pult
0.4 Pult 0.9 Pult
0.5 Pult Pult
8000 6000 4000 2000 0
0
0
16
32
48
64
80
(x) distance from joint (mm) NA–40,2 0.1 Pult 0.2 Pult 0.3 Pult 0.4 Pult 0.6 Pult
0.7 Pult
0.8 Pult
0.9 Pult
0
96
0.1 Pult 0.6 Pult
0.5 Pult Pult
16
32
48
64
80
96
(x) distance from joint (mm) 0.2 PultNA–20,1 0.3 Pult 0.4 Pult 0.5 Pult 0.7 Pult
0.8 Pult
0.9 Pult
Pult
12000
Axial strain (με)
12000
Axial strain (με)
0.2 Pult 0.7 Pult
10000 8000 6000 4000 2000 0
10000 8000 6000 4000 2000 0
0
16
32
48
64
80
(x) distance from joint (mm) NA+0,1
96
0
16
32
48
64
80
(x) distance from joint (mm) NA+20,2
96
Figure
0.1 Pult 0.6 Pult
0.3 Pult 0.8 Pult
0.4 Pult 0.9 Pult
0.5 Pult Pult
10000 8000 6000 4000 2000
0.1 Pult 0.6 Pult
12000
Axial strain (με)
Axial strain (με)
12000
0.2 Pult 0.7 Pult
0.3 Pult 0.8 Pult
0.4 Pult 0.9 Pult
0.5 Pult Pult
10000 8000
6000 4000
2000 0
0 0
16
32
48
64
80
0
96
(x) distance from joint (mm) 0.1 Pult 0.6 Pult
12000
HA–40,1 0.2 Pult 0.3 Pult 0.7 Pult
0.8 Pult
0.4 Pult 0.9 Pult
16
32
48
64
80
96
(x) distance from joint (mm) 0.5 Pult Pult
0.1 Pult 0.6 Pult
HA–20,1 0.2 Pult 0.7 Pult
0.3 Pult 0.8 Pult
0.4 Pult 0.9 Pult
0.5 Pult Pult
12000
Axial strain (με)
Axial strain (με)
0.2 Pult 0.7 Pult
10000 8000 6000 4000
2000 0
10000 8000 6000 4000
2000 0
0
16
32
48
64
80
(x) distance from joint (mm) HA+0,2
96
0
16
32
48
64
80
(x) distance from joint (mm) HA+20,1
96
Figure
0.2 Pult 0.7 Pult
0.3 Pult 0.8 Pult
0.4 Pult 0.9 Pult
0.1 Pult 0.6 Pult
0.5 Pult Pult
40
Lap–shear stress (MPa)
Lap–shear stress (MPa)
0.1 Pult 0.6 Pult
30
20
10
0.3 Pult 0.8 Pult
0.4 Pult 0.9 Pult
0.5 Pult Pult
40
30
20
10
0
0 0
16
32
48
64
80
0
96
0.1 Pult 0.6 Pult
0.2 Pult 0.7 Pult
0.3 Pult 0.8 Pult
0.4 Pult 0.9 Pult
0.1 Pult 0.6 Pult
0.5 Pult Pult
Lap–shear stress (MPa)
40
30
20
10
0 0
16
32
48
64
80
(x) distance from joint (mm) NA+0,1
16
32
48
64
80
96
(x) distance from joint (mm) NA–20,1
(x) distance from joint (mm) NA–40,2
Lap–shear stress (MPa)
0.2 Pult 0.7 Pult
96
0.2 Pult 0.7 Pult
0.3 Pult 0.8 Pult
0.4 Pult 0.9 Pult
0.5 Pult Pult
40
30
20
10
0 0
16
32
48
64
80
(x) distance from joint (mm) NA+20,2
96
Figure
0.6 Pult
0.2 Pult
0.3 Pult
0.7 Pult
0.8 Pult
0.4 Pult 0.9 Pult
0.5 Pult Pult
40 30 20 10 0 0
16
32
48
64
80
96
Lap-shear stress (MPa)
Lap-shear stress (MPa)
0.1 Pult
HA–40,1 0.2 Pult 0.3 Pult
0.4 Pult
0.5 Pult
0.6 Pult
0.7 Pult
0.9 Pult
Pult
30 20 10 0 32
48
64
80
(x) distance from joint (mm) HA+0,2
0.4 Pult
0.5 Pult
0.8 Pult
0.9 Pult
Pult
30
20 10 0
0
16
32
48
64
80
96
96
HA–20,1 Lap-shear stress (MPa)
Lap-shear stress (MPa)
0.1 Pult
16
0.3 Pult
0.7 Pult
(x) distance from joint (mm)
40
0
0.2 Pult
0.6 Pult
40
(x) distance from joint (mm)
0.8 Pult
0.1 Pult
0.1 Pult
0.2 Pult
0.3 Pult
0.4 Pult
0.5 Pult
0.6 Pult
0.7 Pult
0.8 Pult
0.9 Pult
Pult
40 30
20 10 0 0
16
32
48
64
80
(x) distance from joint (mm)
HA+20,1
96