On factors affecting CFRP-steel bonded joints

Construction and Building Materials 226 (2019) 360–375

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On factors affecting CFRP-steel bonded joints Manuel A.G. Silva a,⇑, Hugo Biscaia b, Pedro Ribeiro a a b

Universidade Nova de Lisboa (FCT), Civil Engineering, Caparica, Portugal FSE, UNIDEMI, Universidade Nova de Lisboa (FCT), Caparica, Portugal

h i g h l i g h t s
Durability of CFRP-steel joints under environmental action.
Influence of surface steel treatment for adhesion CFRP-steel.
Reduction of capacity of double strap CFRP-steel joints with temperature.
Influence of loading on bond slip of CFRP-steel joints.

a r t i c l e

i n f o

Article history: Received 18 February 2019 Received in revised form 7 May 2019 Accepted 28 June 2019

Keywords: Surface treatment Adhesion Environmental degradation CFRP-steel joints Thermal effects Bond-slip

a b s t r a c t Failure of structural steel members strengthened with Carbon Fibre Reinforced Polymers (CFRP) may occur at the joints CFRP-steel and this study examines variables that alter or explain the corresponding reduction of load capacity for a specific CFRP laminate, adhesive and steel. Factors and parameters likely to be influential like surface treatment prior to bonding, the bonded length, the glass transition temperature (Tg) of the adhesive, the exposure to aggressive environment, the temperature at service and different types of loading were examined. The experimental program selected double strap CFRP-steel bonded joints under shear for the analysis. The steel surfaces to be bonded were subjected to sand blasting (6.3 bar) or abrasive grinding (6.9 bar) corresponding to thorough blast cleaning Sa2; surfaces rusted after exposure to salt fog at 35 °C were also considered. Differences detected in responses of specimens treated by sand or steel spheres blasting were relatively minor. Tests made at increasing ambient temperatures confirmed that service temperature near and above adhesive Tg caused rapid deterioration of ultimate capacity and change of failure modes. Salt fog cycles (SF) originated the most significant losses of joint capacity. Application of cyclic static loading above the critical loading threshold obtained for unaged joints did not reduce the capacity of joints previously aged by freeze-thaw. The same cyclic loading after salt fog cycles, reduced bond capacity and increase the ultimate slip, suggesting larger effective length. Despite the losses of capacity, microscopic changes of structural nature could not be identified. Ó 2019 Elsevier Ltd. All rights reserved.

1. Introduction Many infrastructures require structural strengthening due to known causes from change of scope of utilization to degradation along their predicted lifetime or design pitfalls. The use of fiber reinforced polymers (FRP) on that endeavor has steadily increased given their high strength to weight ratios, excellent resistance to corrosion and to environmental degradation [1] mainly in reinforced concrete structures, in spite of uncertainties related to their lifetime behavior [2]. Steel structures, namely offshore platforms, tall buildings and steel towers used for antennas as well as steel bridges built a ⇑ Corresponding author at: Faculdade de Ciências e Tecnologia, Universidade Nova de Lisboa (FCT-UNL), 2829-516 Caparica, Portugal. E-mail addresses: [email protected] (M.A.G. Silva), [email protected] (H. Biscaia). https://doi.org/10.1016/j.conbuildmat.2019.06.220 0950-0618/Ó 2019 Elsevier Ltd. All rights reserved.

few decades ago, require rehabilitation worldwide. The substantial market that exists in the area of bridges rehabilitation is easily understood noticing that more than 300,000 railway bridges exist in Europe alone, 22% of steel beam type. In addition, in 2007, 35% of those bridges were more than 100 years old and only 11% had been on service for less than 10 years. In the USA, the Department of Transportation [3] estimated that almost 68.5% of 604,493 bridges were 26 years old or older in 2010, and 11.5% of that total were structurally deficient. At that time, 6 years ago, 200,000 bridges, i.e. one third of the total, were built with steel. The use of CFRP to strengthen metallic structures has been object of many recent research studies e.g. [4–7], and state of the art articles as in [8,9]. Shaat et al. [8] covered publications on retrofit of steel girders, fatigue, durability, surface preparation and applications. The paper by Zhao and Zhang [9], more recent,

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complemented the review concentrating in the bond between steel and FRP, strengthening of steel hollow section members, and fatigue crack propagation. In Japan the externally bonded CFRP laminates and sheets have been applied more and more [10] to rehabilitate metallic structures, weakened by corrosion of tension flanges, fatigue loading and damage or web crippling, a trend also elsewhere in Australia and Asia [11]. FRP sheets and strips adequately bonded to steel structures have shown to improve their mechanical characteristics and fatigue life at cost effective retrofitting techniques for structures not designed to resist seismic action in regions where seismic risk led to new design code provision [12]. Applications of the technique are well documented, e.g. [13–17]. Premature debonding of external strengthening, especially of bonded joints, remains, however, a major source of concern and its likelihood to be influenced by the following parameters and conditions: (i) surface preparation; (ii) adhesive thickness; (iii) overlap length; (iv) geometry of joints; (v) environmental action; (vi) type of adherends; (vii) adhesive curing; (viii) temperature of operation; or (ix) loading conditions. These aspects continue to deserve attention from designers of different types of CFRP-steel structures, e.g. [18–26], whether considered in an isolated form or partially combined. The authors are part of a research team that has been active in studies of adhesively bonded joints of CFRP to steel and literature on their work is available e.g. [27–31]. Given the cited literature and their accumulated experience, the authors compiled obtained results to show the influence of common factors on bond between CFRP and steel joints as discussed in the next subsections. The scope of the work is to summarize and correlate results and provide some guidance on expected behavior and durability of the

joints under stated conditions. The conclusions are naturally limited to cases that fall within the range of described materials and conditions with cautious generalizations conditioned by engineering judgement. 2. Experimental program The performance of adhesively bonded joints is influenced by several factors and, in the case of CFRP-steel, attention is given next to surface treatment, overlap bonded length, environmental aging and operational temperature focusing on their relevance on joint capacity and aspects of bond-slip. The identification of all the tested specimens can be found in Table 1. The geometry, the adhesive and the adherend materials considered in this work are described below. The hereafter designated freeze-thaw cycles had a period of 12 h, imposing an extreme constant value of 20 °C followed by +20 °C, linked by ramps that avoided thermal shock as shown in Fig. 1. The relative humidity (RH) in the chamber was kept at approximately 50% and the cycles were applied without submerging the specimens in water, differently from other studies e.g. [32]. Bonded joints were modeled by double strap CFRP-steel specimens. The CFRP laminates were produced at the Harbin Institute of Technology (HIT) with a thickness of tf = 1.46 mm. Based on three tested flat coupons, their average ultimate tensile stress was 1820 MPa, Young modulus Ef = 180.5 GPa and mean extension 1.00% at failure load. The bonding agent was a two-part room-temperature adhesive (SIKADUR-30) which, according to supplier, had a linear thermal expansion coefficient of 2.5  105/°C, a glass transition temperature 62 °C after 7 days curing at 45 °C, a tensile elastic modulus

Table 1 Designation of the specimens and conditions.

1

ID of the specimens

Bonded length Lb (mm)

Conditions of the specimens

U-T20-0 h-M-01/03 SB-T20-0 h-M-01/03 R-T20-0 h-M-01/03 StB-T20-0 h-M-b10-01/03 StB-T20-0 h-M-b10-04/06 StB-T20-0 h-M-b10-07/09 StB-T20-0 h-M-b10-10/12 StB-T20-0 h-M-b10-13/15 StB-T20-0 h-M-b10-16/18 StB-T20-0 h-M-b10-19/21 StB-T20-0 h-M-b10-22/24 StB-T20-0 h-M-b10-25/27 StB-T20-0 h-M-b10-28/29 StB-T20-0 h-M-b10-30/32 StB-T20-0 h-M-b10-33/35 StB-T20-0 h-M-01/03 StB-T20-0 h-M-01/03-a1 StB-T20-0 h-M-04/06 StB-T35-0 h-M-01/02 StB-T50-0 h-M-01/02 StB-T65-0 h-M-01/02 StB-T80-0 h-M-01/02 StB-T95-0 h-M-01/02 StB-T20-FT1000 h-M-01/03 StB-T20-FT2500 h-M-01/03 StB-T20-FT3500 h-M-01/03 StB-T20-FT5000 h-M-01/03 StB-T20-SF5000 h-M-01/03-a1 StB-T20-SF5000 h-C-01/03-a1 StB-T20-0 h-C-01/03 StB-T20-FT1000 h-C-01/03 StB-T20-FT2500 h-C-01/03 StB-T20-FT3500 h-C-01

200

Untreated, 20 °C, 0 h, monotonic test Sandblasted, 20 °C, 0 h, monotonic test Rusted (after SF exposure), 20 °C, 0 h, monotonic test Steel blasted, 20 °C, 0 h, monotonic test

10 25 50 60 75 90 100 110 125 135 150 200 150 200

The resin used in these specimens was from a different pot.

Steel Steel Steel Steel Steel Steel Steel Steel Steel Steel Steel Steel Steel Steel Steel

blasted, blasted, blasted, blasted, blasted, blasted, blasted, blasted, blasted, blasted, blasted, blasted, blasted, blasted, blasted,

35 °C, 50 °C, 65 °C, 80 °C, 95 °C, 20 °C, 20 °C, 20 °C, 20 °C, 20 °C, 20 °C, 20 °C, 20 °C, 20 °C, 20 °C,

0 h, monotonic test 0 h, monotonic test 0 h, monotonic test 0 h, monotonic test 0 h, monotonic test 1000 h, monotonic test 2500 h, monotonic test 3500 h, monotonic test 5000 h, monotonic test 5000 h, salt fog, monotonic test 5000 h, salt fog, cyclic test 0 h, cyclic test 1000 h, freeze–thaw, cyclic test 2500 h, freeze–thaw, cyclic test 3500 h, freeze–thaw, cyclic test

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Fig. 1. Graphical representation of 12 h freeze-thaw temperature cycle (RH = 50%).

of 11.2 GPa and an average bonding stress to steel above 21 MPa for a steel surface sandblasted to degree Sa2.5 with siliceous sand passing sieve number 4 (up to 4.76 mm) and pressure 6.3 bar. Nine dog bone coupons of the adhesive were aged and tested, at 0 h and after 1000 h and 2500 h of FT cycles, as reported in a subsequent section. The clamped extremity of the specimens with bf = 20 mm had a bond length of Lb = 90 mm while at the other end it was either 200 mm or 150 mm depending on whether or not the specimens were subjected to temperature increases. The value of 150 mm was based on an estimate of the effective length (Leff) as in [30]. Specimens were kept 48 h at 20 °C after bonding, and then placed in a climatic chamber at 35 °C and 50% relative humidity for 7 days to accelerate the cure of the adhesive. The surface of the HIT CFRP laminates was cleaned with acetone. The joints were assembled by bringing two aligned steel plates within a small distance and bonding two rectangular strips of CFRP, each one to one face of the steel plates, with adhesive SIKADUR-30. The adhesive thickness was kept at 1 mm with steel spacers. The steel plates were ts = 5 mm thick, averaged yield stress

382.5 MPa, ultimate stress 560.5 MPa, strain at ultimate stress 14.5% and elasticity modulus Es = 179 GPa. Due to the number of tests carried out, two different pots of resin were required, and specimens assembled from the second resin mix are identified in Table 1 with a letter ‘‘a” at the end of their ID label as shown in Table 1. The illustrations in Fig. 2 correspond to the cases of overlap bonded length Lb = 150 mm and CFRP width bf = 20 mm. Tests were performed with a Zwick 50 kN tensile machine and a tensile force rate corresponding to a displacement of 1 mm/min, except if otherwise stated. At this rate the capture of points for the bond-slip curves, especially on the softening phase, could be done with enough detail. The importance of the treatment of the steel surface, as well as the influence of operating temperature and thermal cycles around 0 °C on the shear capacity of double-strap CFRP-steel joints for the component materials utilized in the present work were experimentally examined in terms of joint capacity and bond-slip curves and are reported next.

3. Surface preparation The roughness of either CFRP or steel surfaces relates to their texture and influences the bond capacity of the joints. It is also a characteristic much affected by surface treatment and a study was made on the influence of the roughness of the steel surface. For a relatively smooth surface, increasing the roughness makes larger the contact area available to the adhesive larger, the surface energy higher and even a moderate increase determines an improvement in adhesion. If roughness is larger, its effect on wetting will also be larger, but significant roughness can reduce wetting achieved at equilibrium, thereby reducing, adsorption, and fundamental adhesion [33]. There is an optimum surface roughness, leading to a peak of strength, i.e. excessive surface treatment may be counterproductive [34]. Adherend surface roughness leads to maximum adhesive bond strength when it varies in the range of 1.5–2.5 lm, and to lower shear stresses both for very smooth surfaces, with mean roughness Ra < 1.0 lm, and very rough surfaces,

Fig. 2. (a) Test in Zwick machine; (b) geometry of double strap steel/CFRP specimens; and (c) aspect of some double strap specimens.

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Ra > 2.5 lm [34], Ra being a current parameter that measures the average mean deviation about the centre line (or surface in 3-D analysis). The shear capacity of the joints is also related to interlocking and increased bonding area and to the chemical bond between the adhesive and adherend [35]. Some earlier data on this topic were studied by Teng et al. [36] who considered five different surface preparation methods and different adhesives and focused on conditions that favor cohesive instead of adhesive failure of the joints. They observed that the surface energy of the steel should exceed a certain threshold value (50 mJ/m2) and that grit residues from blasting alter the chemical composition of the steel surface and should be compatible with the adhesive, recommending aluminum grit. In the present study, the steel surfaces were either subjected to sand blasting (6.3 bar) as described above, or to grinding with steel spheres diameter 1.8 mm to level Sa 2 (6.9 bar). Exposure of untreated surfaces to 250 cycles of salt fog for 250 h at 35 °C (Fig. 3) was also imposed to rust the surface. Table 2 shows the values of roughness parameters and degree of roughness. Three dimensional amplitude parameters were selected to characterize the topography roughness in terms of arithmetic mean deviation PSa and root mean square deviation PSq (Msystem) [37–39], no filters. The effects of surface topography on shear strength of lap adhesive joints of steel as well as the influence of the chemical bond characteristics of the adhesive and adherend materials have been studied [40,41]. Fig. 4 displays images of surfaces prior to surface preparation and after sand blasting and steel grinding, amplified 100 and 200 times. The mechanical tests run to compare different performances due to the applied surface treatment imposed a displacement rate higher than in the remaining work, approximately 15 mm/min. Table 3 displays average maximum bond stress, maximum and ultimate slip, and (half) shear capacity of the double strap joints, found experimentally with the described treatments and materials. For the conditions described, blasting with small spheres led to slightly higher capacity, with least scattering, while no treatment was the worst option. The capacity obtained for surfaces rusted by salt fog are misleading in that more prolonged and detailed salt fogging caused severe loss of capacity reported elsewhere. Fig. 5 shows the local surface roughness of two samples extracted after treatment by sand blasting and by steel spheres. The images were obtained with a confocal microscope Zeiss LSM 700. Fig. 6 is divided in two parts and shows load-slip curves on the left-hand side, and local bond-slip curves on the right-hand side for surfaces with different roughness corresponding to (a) no surface preparation; (b) sand blasting; (c) steel gritting; (d) after salt fog corrosion. The highest average maximum bond stresses were attained for blasting with small spheres or sand, and the largest capacity of the specimens was 2F = 33.4 kN for steel spheres grind-

Table 2 Surface topography for each condition of the surfaces. Surface preparation

PSq (lm)

PSa (lm)

Degree of roughness (ISSO 1302)

Untreated Sand blasted Steel blasted Rusty

3.14 8.95 22.48 33.64

2.43 7.23 19.25 25.50

N7 N9 N10 N11

ing. The descending branches of the curves led to very small differences between sult and smax. The specimens treated by steel gritting had the smallest average smax = 0.06 mm and the average ultimate slip sult = 0.13 mm. All the reference joints were made with HIT laminates. The tests showed essentially cohesive failures, Fig. 7, unlike in some other research programs with different geometries and adhesives [42] or at higher temperatures, where adhesive failure modes were reported [31]. The influence of the adhesive was e.g. shown in [43] where specimens bonded with SIKA 30 confirmed the dominant cohesive failure found in the present study, whereas adhesive Sika 330 led to interface debonding, mostly between CFRP and adhesive. The authors also showed that specimens without steel surface treatment failed by interface debonding, especially between the steel and the adhesive. The summarized results evidence the need for more tests and consideration of additional degrees of roughness and type of treatment of the steel surface to find values of optimum treatment defined by maximum achieved capacity of the joints under tensile loading. Such future study ought to include other parameters like CFRP surface wettability and rugosity, and more than one type of CFRP laminate and adhesive. 4. Bonded length Several specimens, with different bonded lengths and either 10 mm or 20 mm width, were tested to study the influence of the bonded length on the bond capacity of the CFRP-steel interfaces. The maximum loads transmitted to the CFRP laminate are summarized in Table 4. As expected, they increased with the overlap bonded length till a maximum was reached, staying constant beyond a certain threshold. The study of effective length was the subject of several publications associated with retrofitting of structures e.g. [44–48]. Following a methodology based on [45], the estimated effective bond length of the CFRP-steel interface was Leff = 136 mm, exceeded by the actual bonded lengths 150 and 200 mm of the tested specimens. The predicted maximum load of the CFRP-steel interface (for bf = 10 mm) was 6.75 kN. Fig. 8 shows how the estimations fit the experimental data that define the shaded area. The way the scattering of experimental results for Fmax was affected by

Fig. 3. Steel surface with: (a) no surface preparation; (b) sand blasted; (c) steel blasted; and (d) salt fog exposition.

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(a)

(b)

No treatment

Sand blasted

Steel spheres blast

Fig. 4. Coupled images of steel surfaces, amplification: (a) 200; and (b) 100 below.

Table 3 Bond maximum stress, slips, half joint capacity and strain for different steel surface preparations. Surface

Untreated Sand blasted Steel blasted Rusted

rmax (MPa)

smax (mm)

sult (mm)

emax (gauge)

Fmax (kN)

Avg

Stdev

Avg

Stdev

Avg

Stdev

Avg

Stdev

Avg

Stdev

19.36 22.93 26.07 18.78

4.72 4.17 1.22 2.36

0.07 0.07 0.06 0.11

0.01 0.02 0.00 0.08

0.08 0.14 0.13 0.16

0.01 0.08 0.12 0.05

15.8 15.4 16.7 17.4

1.0 0.7 0.0 0.7

0.30 0.29 0.30 0.33

0.02 0.03 0.02 0.01

Stdev = standard deviation; Avg = average.

Fig. 5. 3-D surface roughness PSq after sand blasting and treatment with small spheres, respectively, 8.95 lm and 22.48 lm.

the bonded lengths is also indicated by the listed coefficients of variation (CoV) in Table 4. 5. Increase of temperature Temperature may change the behaviour of the adhesive in a significant way when it nears the glass transition value Tg and lowers the capacity of transmission of the loading between the bonded materials [49]. Available modelling exists for CFRP/concrete joints that introduces the dependency of material properties on the temperature namely one by Dai et al. [50]. The authors present a non-

linear local bond-slip model for elevated temperature as a generalization of a two-parameter model for ambient temperature. Another model, also for joints of CFRP and concrete, was proposed by Dong and Hu [51] where a relation of bond properties with the temperature difference (T  Tg) is submitted. Li et al. [52] adjusted the model described in [50] to steel substrates and, more recently [31], proposed a temperature-dependent bond-slip model for CFRP-steel interfaces. The effects that environment temperature had on strength and ultimate strains of the CFRP-steel joints loaded at temperatures above that at their initial assemblage, or approaching the Tg of

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Fig. 6. Load-slip (on the left side) and local bond-slip (on the right side) curves corresponding to different surfaces: (a) with no surface preparation; (b) sand blasted; (c) steel blasted; (d) after salt fog corrosion.

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Fig. 7. Cohesive failure surfaces of reference specimens, after monotonic loading.

Table 4 Maximum loads of the double strap bonded joints. ID of the specimens

Lb (mm)

Fmax (kN)

Stdev (kN)

CoV

StB-T20-0 h-M-b10-01/03 StB-T20-0 h-M-b10-04/06 StB-T20-0 h-M-b10-07/09 StB-T20-0 h-M-b10-10/12 StB-T20-0 h-M-b10-13/15 StB-T20-0 h-M-b10-16/18 StB-T20-0 h-M-b10-19/21 StB-T20-0 h-M-b10-22/24 StB-T20-0 h-M-b10-25/27 StB-T20-0 h-M-b10-28/29 StB-T20-0 h-M-b10-30/32 StB-T20-0 h-M-b10-33/35

10 25 50 60 75 90 100 110 125 135 150 200

0.96 2.82 5.59 6.07 5.18 7.16 5.66 6.25 6.10 7.12 5.84 6.59

0.56 0.42 1.44 1.23 0.34 0.23 0.51 0.74 1.10 0.15 1.24 0.76

0.56 0.15 0.26 0.20 0.07 0.03 0.09 0.12 0.18 0.02 0.21 0.12

Stdev = standard deviation; CoV = Coefficient of Variation.

the load transmitted to each CFRP strip. The slips including the contribution of the thermal dilation of the materials were obtained [53]:


ds ¼ ð1 þ bÞ  ef þ b  af þ as  DT dx

ð1Þ

where ef is the strain in the CFRP laminate, af and as are, respectively, the thermal expansion coefficients of the CFRP laminate and of the steel bar, b is the axial stiffness ratio defined according to

b¼

Ef t f bf Es t s bs

ð2Þ

and

DT ¼ T  T 0 Fig. 8. Maximum loads transmitted to the CFRP laminate with different overlap bonded lengths.

the adhesive, are described below. The double lap shear joints with 20 mm wide CFRP strips were tested in a Zwick machine with a heating chamber. The temperature of the specimens never reached the Tg of the resin composing the matrix (approximately 103 °C as determined by DMA). The average results are summarized in Table 5. The failure loads decreased with the temperature. The average ultimate strains also decayed, especially for temperatures near and above Tg of the adhesive (around 67 °C as mentioned earlier). The experimental tests led to curves relating applied traction loads and slip as shown in the load-slip curve in Fig. 9 where F is

ð3Þ

where T0 is the temperature prior to development of inner strains and T the stabilized oven temperature when the mechanical load is applied to the specimens. The bond-slip data for the 20, 35, 50, 65 and 80 °C led to the local bond-slip curve in Fig. 9. Numerically those results were approximated by a bi-linear shape defined by the bond stress peak, or maximum bond stress (smax), the maximum slip (smax) meaning the slip developed at smax, and the ultimate slip (sult) meaning the slip beyond which no more load transfer takes place. Observing the average values, the increase of the temperature did not affect much the values of the maximum slip, and, in the softening stage, for temperatures higher than Tg of the adhesive, the values of bond tended to vanish. In general, the adhesively bonded joints at higher temperatures developed a much lower peak bond stress. Comparing the average

Table 5 Decrease of capacity of joints with temperature.

1

Temperature (°C)

20

35

50

65

80

95

Average load capacity (kN) Average ultimate strain %

33.3 0.32

32.7 0.29

29.7 0.231

25.0 0.28

12.2 0.10

11.4 0.12

One single test available.

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Fig. 9. Curves of load-slip on the left side and local bond-slip on the right side for (a) 20 °C and 35 °C; (b) 50 °C; (c) 65 °C; and (d) 80 °C.

results at 20 °C and 65 °C, the maximum bond stress decreased approximately 75% and the corresponding slip increased from 0.082 mm to 0.220 mm (an increase of 171%).

Regarding the results obtained for ultimate slip it increased 60% from 0.20 mm at 20 °C to 0.32 mm at 65 °C showing a large increment of the relative displacement of CFRP and steel. Results

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Fig. 10. Failure surfaces at: (a) 20 °C; (b) 35 °C; (c) 65 °C; and (d) 80 °C.

obtained at 95 °C, a temperature higher than the Tg of the adhesive, are not displayed here, but showed a maximum bond stress of only 3.90 MPa. Fig. 10 shows images of the failure surfaces of specimens tested at 20, 35, 65 and 80 °C. Up to 50 °C the failures were cohesive in the adhesive, starting near the loaded end, followed by debonding of the entire strip. At temperatures close to Tg of the SIKADUR-30, the adhesive mode failure appeared, even though at 65 °C there were signs of both adhesive and cohesive failure, leading to designating it as a hybrid mode. At 80 °C the failure was essentially adhesive at the interface. The initial stiffness of the specimens was also reduced for higher temperatures, as expected from the tensile tests of the adhesive showing a decrease of the interfacial stiffness of the interface. Therefore, the interfacial elastic stage of the CFRTP-steel interface was severely affected by the temperature increase. 6. Freeze-thaw aging Freeze-thaw cycles were not expected to cause any chemical changes, and physical changes were anticipated only in cases when different materials forced to expand differently might end up forming microcracks resulting from the induced stress field. The variation of Tg of the adhesive, due to FT cycles, was nonetheless obtained by DMA with a model Q800 from TA Instruments and obtained results shown in Table 6. The bone shaped adhesive coupons fractured when subjected to tensile tests and the average values of the ultimate tensile deformation and stress appear in Table 7, for reference coupons and

Table 6 Tg for SIKADUR-30 after FT cycles. Time of exposure (h)

Tg (°C)

0 2500 5000

67.0 66.8 64.5

coupons subjected to 1000 h and 2500 h of FT cycles, showing an increase of 17.8% of the ultimate strain and a 14.5% decrease of the ultimate stress, from 0 h to 2500 h. Fig. 11 shows images of the failure surface of tensile tested coupons of adhesive SIKADUR-30, obtained with a SEM-Hitachi TM 3030Plus Tabletop, amplified 300 and 600 after 2500 h of FT cycling. Some cracking seems apparent in the latter image. Small holes seen on the adhesive correspond to air bubbles entrapped at the time of preparation of the adhesive mixture. The heterogeneity of the observed failure surfaces made their comparative characterization difficult. Differences are hard to find, and a much longer period is thought to be required for the defined FT to cause noticeable changes. This is also due to the absence in these adhesive coupons of constraints susceptible to create internal states of strain upon temperature rise. The average percent weights of the chemical constituent parts found in the adhesive at 0 h, 1000 h and 2500 h were examined, but the changes found were attributed to localized differences of the spot sampling. The tests of 12 joint specimens led to results at 0 h, 1000 h, 2500 h and 3500 h. The differences found were negligible and Table 8 shows the values obtained indicating an increase of the average ultimate slip, with the other quantities presenting very small oscillations. The failure modes were classified by spotting attachments of resin on the debonded surface of the CFRP laminate and/or on the corresponding surface of the steel plates. The cohesive failure mode started for these joints in the adhesive adjacent to the steel rather than to the CFRP laminates. Fig. 12a) and b) correspond to images of failure surfaces for 0 h and after 3500 h of cycles [20, +20]°C, and correspond to an area below the first strain gauge, where the CFRP surface after debonding showed very few and small traces of adhesive. Fig. 12c shows a different area of typical cohesive failure with the surface of the laminate almost totally covered with chunks of adhesive, after 5000 h of freeze-thaw cycles. As shown in the latter part of the experimental program, after 10,000 h the failure surface on the interface with steel had almost no adhesive since failure occurred at the interface adhesive/steel

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M.A.G. Silva et al. / Construction and Building Materials 226 (2019) 360–375 Table 7 Adhesive tensile ultimate stress and deformation. Mechanical properties

Reference

1000 h

2500 h

Coupons

Average

Coupons

Average

Coupons

Average

Ultimate stress (MPa)

21.91 23.11 26.78 24.31

24.03

20.33 29.06 21.86 23.01

23.57

23.81 21.49 14.66 22.19

20.54

Ultimate strain (%)

0.283 0.317 0.392 0.308

0.325

0.459 0.376 0.391 0.436

0.415

0.342 0.353 0.383 0.455

0.383

(a) 300×

(b) 600×

Fig. 11. Failure surfaces after tensile tests of adhesive aged by cycles of freeze-thaw for: (a) 1000 h (amplification: 300); and (b) 2500 h of FT (amplification: 600).

Table 8 Bond stresses, interfacial slips and maximum loads after freeze-thaw cycles. Aging (h)

rmax (Mpa) Avg

Stdev

Avg

Stdev

Avg

Stdev

Avg

Stdev

Max

0

19.1 20.0 19.8

19.6

0.46

0.07 0.08 0.08

0.08

0.01

0.24 0.23 0.32

0.26

0.05

16.2 17.9 16.9

17.0

0.8

17.9

1000

21.7 23.6 27.3

24.2

2.84

0.09 0.06 0.06

0.07

0.02

0.38 0.24 0.29

0.30

0.07

19.0 16.3 16.4

17.2

1.5

19.0

2500

27.0 22.3 21.2

23.5

3.11

0.08 0.07 0.08

0.07

0.00

0.33 0.22 0.28

0.28

0.05

17.9 18.6 18.2

18.2

0.3

18.6

3500

22.7 26.7 18.5

22.7

4.09

0.07 0.09 0.08

0.08

0.01

0.27 0.30 0.34

0.30

0.04

17.0 16.9 16.7

16.8

0.1

17.0

smax (mm)

sult (mm)

Fmax (kN)

Avg = average; Stdev = standard deviation.

[54]. It is also reported in the latter study, for different CFRP laminates and the same cycles [20, +20] °C, that the failure mode was mostly adhesive because the bond CFRP-adhesive was not strong enough to transfer load capable of rupturing the adhesive itself. Further studies are necessary with a considerably larger number of applied cycles, also of larger amplitude, to verify the magnitude of the reduction of carrying capacity and relate it to design requirements. The associated effects of i) diffusion of larger masses of moisture, and ii) more micro-cracking and further degradation due to freezing recommend the use of immersion on future FT tests. Fig. 13 shows the local bond slip curves of CFRP-to-steel interface at 0 h of exposure (reference) and after 3500 h exposure to

freeze-thaw cycles. The results showed that the bond-slip relationship could be essentially described by a bi-linear relationship. Despite some discrepancies, the average values of those responses did not show relevant changes among them. For instance, the maximum bond stress after 3500 h of exposure led to an increase of 17.4% (from 17.0 MPa to 16.8 MPa) whereas the corresponding maximum slips varied only from 0.082 mm at 0 h to 0.084 mm after 3500 h of ageing exposure, differences that were mostly negligible given the inherent scatter of those tests. 7. Salt fog cycles A corrosion box was used to subject the joints to accelerated degradation by salt fog using a solution with 5% weight of NaCl

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Fig. 12. Failure surfaces for: (a) 0 h; (b) 3500 h, a laminate region, amplified 200; (c) 5000 h, adhesive on laminate, 265; and (d) adhesive on steel, 600.

per litre of distilled water and generating 6 h of spray followed by 18 h ‘‘drying” at a constant temperature of 35 °C. The specimens subjected to salt fogging for 5000 h were tested to compare their performance with the results obtained at 0 h. The effects on steel were severe as seen in Fig. 14 where the load-slip and local bond-slip curves after 5000 h of SF cycles are displayed and can be compared with those obtained at 0 h that define the gray areas. The comparisons show a considerably lower ultimate load, a reduced maximum bond stress and also a smaller sult. These effects were caused by the lowering of the bond strength at the interface between CFRP and adhesive and also by the corrosion of part of the steel bonded to the laminates. Fig. 15 illustrates the typical rusted aspect of the specimens after 5000 h of SF exposure (Fig. 15c-d) compared with reference specimens (Fig. 15a-b). The corrosion penetrated within the bonded area affecting the bond between the adhesive and the steel as seen in Fig. 15d. Some specimens left inside the SF chamber for an exposure up to 10,000 h and all debonded while in the chamber.

8. Loading history – cyclic tests pre and after aging Double strap shear joints were subjected to cyclic shear loading, both unaged and, later, after freeze-thaw and salt fog aging to verify possible accumulation of effects. The loading cycles were applied without inducing dynamic effects and therefore designated as a quasi-static (pseudo) cyclic loading. A ‘‘threshold loading” below which the quasi-static cycles caused no damage was tentatively found at approximately 50% of the monotonic failure load. Above that value, damage was localized near the CFRP loaded end [28]. The loading/unloading tests protocol programmed three cycles with the same amplitude, succes-

sively followed by sets of three cycles with a larger peak. The first three load/unload cycles subjected the CFRP-steel interface to 25% of the ultimate load reached in the monotonic tests. In the following three cycles the peak load would reach 50% of the ultimate monotonic load, proceeding in 10% peak load increments. The loading stopped when failure took place or at the end of the 18 pseudo-cyclic loading defined in Fig. 16. The presence of eventual cycles with amplitude above the maximum monotonic load was considered (see Fig. 16) to ensure the failure of the specimens, should any of them reach such level of strength, situation that did not take place in the actual tests. 8.1. After freeze-thaw aging The curves of applied load and bond stress versus slip are displayed in Fig. 17 for the aged specimens after exposed to 1000 h, 2500 h and 35,000 h of FT cycles. At each stage the bond-slip curves of the reference specimens (in shaded areas) were compared with the after aging response curves for cyclic loading. None of the FT aged specimens failed for a load showing relevant changes when compared with the reference curves, i.e. the cyclic static loading did not damage significantly the specimens. This means that the CFRP-steel interface did not accumulate slips during the unloading phases and no relevant hysteresis loop was observed. Fig. 18 shows how the interfacial slip developed at the CFRP loaded end with the half-cycles (loading or reloading phases) imposed in the tests. Only three specimens reached the 18.5th half-cycle (specimens StB-T20-FT2500h-C-01, StB-T20-FT2500hC-02 and StB-T20-FT3500h-C-01), i.e. the first half-cycle corresponding to 100% of the maximum monotonic load. The interfacial slips increase with the number of half-cycles but with a lower

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371

Fig. 13. Load-slip (on the left side) and local bond-slip (on the right side) curves: (a) after 1000 h of FT cycles; (b) after 2500 h of FT cycles; and (c) after 3500 h of FT cycles.

increase for the same set of three cycles with the same load transmitted to the CFRP laminate. As the peak of the 19th cycle approaches, designated for easier reference as 18.5th peak, the interfacial slips show the highest slip increment. Since the determination of the effective bond length is closely related to the ultimate slip [30], these results suggest that no meaningful change on the Leff of the CFRP-steel interface is expected due to the imposed cyclic loading. 8.2. After salt fog aging The same loading protocol was assumed for the specimens exposed to 5000 h of salt fog. The load-slip and local bond-slip curves obtained from the tested specimens are shown in Fig. 19.

The results were compared with the corresponding monotonic results after 5000 h of SF exposure. Both the load-slip and the local bond-slip envelopes of the three specimens (StB-T20-SF5000h-C01/03-a) followed the corresponding monotonic curves without relevant discrepancies. The interfacial maximum bond stress and slip did not change during the first three loading cycles, but the softening stage was slightly affected by the exposure to SF and the bond stresses decayed more rapidly than for the monotonic tests after the maximum bond stress was exceeded. So, the ultimate slip of the cyclic tests seems to reduce from its initial value, i.e. obtained from the monotonic curves of the aged specimens. The calculated local behaviour of the three specimens showed almost no hysteresis loop and specimen StB-T20-SF5000h-C-03-a

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Fig. 14. Comparison between the reference specimens (at 0 h) and the aged ones after 5000 h of SF cycles: (a) load-slip curves; and (b) local bond-slip curves.

Fig. 15. Aspect of the: (a) reference specimens; (b) failure mode of the reference specimens; (c) specimens after 5000 h of SF exposure; and (d) failure mode of the specimens after 5000 h of SF exposure.

The interfacial slips developed at the CFRP loaded end increased in all the reloading phases of the adopted cyclic loading. However, the rupture of these specimens occurred at a higher half-cycle (21.5th peak) than that registered from the FT specimens (18.5th peak). Fig. 20 shows how the slips at the CFRP loaded end evolve with the cycles imposed to the CFRP-steel aged interface. From this figure, differently from what was previously observed after FT, the ultimate slip was larger than the corresponding reference average value. This suggested, unlike the FT aged specimens, that some change on the effective bond length was induced by the adopted cyclic loading protocol. Previous references to the influence of cyclic loading on the effective bond length are not known, but the results suggested an increase of the effective bond length of CFRP-steel under cyclic loading to be further studied. Fig. 16. Loading protocol followed in the cyclic tests.

9. Conclusions showed some residual slip (measured at x = 12.5 mm from the CFRP loaded end). The maximum loads obtained from the monotonic and from the cyclic tests after 5000 h of SF exposure were quite similar.

Summarizing results of the tests and observations made on the described double strap shear CFRP-steel bonded joints, conclusions were reached that indicate the importance of each of the corresponding parameters:

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373

Fig. 17. Load-slip (on the left side) and local bond-slip (on the right side) curves: (a) after 1000 h of FT cycles; (b) after 2500 h of FT cycles; and (c) after 3500 h of FT cycles.

– Surface treatment of the steel plates improved the shear capacity of the joints. – Steel surface rust after 10 days of exposure to salt fog conduced to circumstantial higher failure shear load due to a layer of rust that resisted to detachment. Further studies on steel surface preparation are necessary to establish the degree of roughness that leads to highest values of shear strength in the interface for each type of treatment. The importance of the roughness and wettability of the surface of the CFRP strips was not investigated but needs research.

– The increase of the overlap bonded lengths increased the bond capacity of the CFRP-steel interfaces till an effective bond length of 136 mm. – Rise of operational temperature caused significant lowering of the carrying capacity of the joints especially when the temperature was near and above the value of the glass transition temperature Tg of the bonding adhesive. Aging by thermal cycles [20, +20]°C (‘‘freeze-thaw”) for 3500 h, caused only minor changes in the glass transition temperature Tg of the adhesive, and on the CFRP-steel joint strength. Microscopic changes of structural nature could not be identified.

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Fig. 20. Interfacial slips developed at each half-cycle (reloading phases) of the specimens after exposure to 5000 h of SF aging.

– Salt fog cycles caused severe degradation of ultimate capacity after 5000 h of exposure. – Specimens aged by ‘‘freeze-thaw” and subjected to quasi-static cyclic loading applied above the ‘‘critical threshold” did not show reduction of capacity. However, the same cyclic loading after exposure to salt fog cycles, reduced bond capacity and the ultimate slip increased, suggesting that the effective bond length might have increased. – SEM techniques proved valuable to identify microdamage and genesis of failures. Coupled EDS determinations of chemical alterations, however, would require a much larger number of analyses unfeasible in the framework of this study.

Declaration of Competing Interest None. Acknowledgements

Fig. 18. Interfacial slips developed at each half-cycle (reloading phases): (a) after 1000 h of FT cycles; (b) after 2500 h of FT cycles; and (c) after 3500 h of FT cycles.

Yongming Yang cooperation with the junior author in some experimental tests is acknowledged. H. Biscaia expresses his personal gratitude to Fundação para a Ciência e Tecnologia (FCT) for his research grant SFRH/BPD/111787/2015 obtained under Project UID/EMS/00667/2019.

Fig. 19. Comparison between monotonic and cyclic loading of unaged and aged specimens (5000 h of SF cycles): (a) load-slip curves; (b) local bond-slip curves.

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