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Impact of corrosion on bond in uncracked concrete with confined and unconfined rebar
Construction and Building Materials 155 (2017) 550–559
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Impact of corrosion on bond in uncracked concrete with confined and unconfined rebar David W. Law ⇑, Thomas C.K. Molyneaux School of Civil Engineering, RMIT University, Melbourne, Australia
h i g h l i g h t s For confined specimens bond strength initially increases as corrosion proceeds. Unconfined specimens display no clear relationship. The variation between top and bottom cast bars reduced as cover increased. The effect of confinement reduced as cover increased. There was little scaling effect between the 12 and 16 mm bars.
a r t i c l e
i n f o
Article history: Received 30 January 2017 Received in revised form 16 August 2017 Accepted 18 August 2017
Keywords: Bond Corrosion Rebar Cover Concrete Cracking
a b s t r a c t This paper examines the variation in bond strength in corroding specimens, before surface cracks are visible. Reinforced concrete specimens containing 16 mm and 12 mm diameter bars at covers of 1, 2 and 3 times bar diameter, with and without confinement, were subjected to accelerated corrosion. Bond strength is seen to increase during the early stages of corrosion (pre external cracking) for confined bars at all covers and unconfined bars at 3 times cover. However, while 16 mm bars, 1 and 2 times cover showed similar behaviour the 12 mm bars displayed a decrease in bond once corrosion was initiated. Ó 2017 Elsevier Ltd. All rights reserved.
1. Introduction Corrosion of reinforcing steel in concrete structures is one of the primary causes of concern for owners with billions of pounds being spent on repair and maintenance. The bond between the steel reinforcement and the concrete is essential for the reinforced concrete to act in a composite manner and the presence of corrosion threatens this integrity. At present much of the inspection and remediation is conducted when there is visual evidence of corrosion, such as rust staining and cracking/spalling. However there is a significant period during which corrosion is continuing and the bond strength changing whilst there are no surface signs. Corrosion of reinforcing steel in concrete is a complex electrochemical process. Reinforcement in concrete is resistant to corrosion due to the protective iron oxide film that arises from
⇑ Corresponding author. E-mail address: [email protected] (D.W. Law). http://dx.doi.org/10.1016/j.conbuildmat.2017.08.112 0950-0618/Ó 2017 Elsevier Ltd. All rights reserved.
the alkaline conditions (a pH of approximately 12.5) provided by the surrounding concrete. Chloride ions however, from an external marine environment or from de-icing salt can impair the protective oxide film and initiate corrosion. Variations in properties at the surface of the bar encourage a mixed electrode to form in which one region of the steel acts as an anode and another as the cathode with pore water and dissolved ions forming an electrolytic cell. At the anode iron atoms pass into solution as ferrous ions (positively charged). For corrosion to occur a loss of passivation, the presence of oxygen and the presence of water are all required. The deterioration of reinforced concrete is characterized by a general or localized loss of section of the reinforcing bars and the formation of expansive corrosion products at the surface of the bar. The production of expansive products creates tensile hoop stresses around the bar within the concrete, which can result in cracking and spalling of the concrete cover. This cracking can lead to accelerated ingress of aggressive agents thus accelerating further corrosion. This process leads to loss of strength and stiffness
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of the concrete cover and threatens the bond strength between the concrete and the reinforcing steel. Finally the corrosion reduces the cross section of the reinforcing steel, which can affect the ductility of the steel and the load bearing capacity, which can ultimately impact upon the serviceability of the structure and the structural capacity [1,2]. During the corrosion process the volume of the corrosion products created on the surface of the bar is greater than that of the steel consumed during the process of corrosion and hence radial and circumferential stresses are generated in the concrete around the steel bars, resulting in the cracking of the surrounding concrete. The mass of steel lost due to a corrosion current (I) can be calculated by applying Faraday’s law.
Dm ¼ MIt=zF
ð1Þ
where Dm is the mass of steel consumed (g), M is the atomic weight of the metal (56 g for iron); t is the time in seconds; z is the ionic charge (2 for iron); and F is Faraday’s constant (96,500 A/s). There has been a significant effort in assessing and modeling the degradation of reinforced concrete structures in order to assist asset managers in adequately planning for the prevention/minimisation of corrosion damage and maintenance, with a number of models proposed to predict the effect of corrosion on bond [2–4]. Research has been reported on the relationship between the corrosion level (corrosion penetration or mass loss of the steel) and cracks [5–14]; current density versus surface crack width [1,15,16]; and bond strength versus corrosion level (mass loss or corrosion penetration). The use of shear reinforcement within the concrete can counteract the loss of bond capacity to a certain degree by providing confinement and resisting the radial tensile cracking. Work by Wang et al. (2004) shows that bond strength increases because of confining action given by stirrups [17–19]. However, relatively little research has studied the relationship between bond strength and surface crack width [16,20–22] with even less on the effect of corrosion on bond prior to cracking [23]. Previous research has shown an initial increase in bond strength for corroded confined when cracking first occurs, however, for unconfined bars a decrease in bond strength is observed with the initial cracking [16,21]. However the influence of the level of corrosion prior to cracking is not clear. This paper reports the effect of corrosion level on confined and unconfined bars of 12 mm and 16 mm diameter, with 1, 2 and 3 times cover up to the initiation of the first crack.
embedded bar, the mean stress at peak pullout force is representative of the relative structural capacity. Specimens of rectangular cross section were cast with a longitudinal reinforcing bar in each corner. Specimens with confined and unconfined 12 mm and 16 mm diameter bars were cast with covers of 1, 2 and 3 times bar diameter. The specimens were dimensioned: 380 mm length, 200 mm width and 300 mm depth with an 80 mm plastic tube placed on the far (unloaded) end of each steel reinforcing bar to create a de-bonded zone which protects the reinforcement from the confining pressure of the concrete due to the transverse reaction (supporting) force during the pullout testing, Fig. 1. Confinement, where applicable, comprised undeformed stainless steel 6 mm diameter links at 100 mm spacing, three links in each specimen, positioned symmetrically between the end of the plastic tube and the end of the specimen. The specimens are similar in form to those adopted by [24]. Four specimens of each type were cast making a total of 48 specimens (3 covers, 2 bar sizes, confined and unconfined). The longitudinal reinforcement selected comprised standard deformed bars of 12 mm and 16 mm diameter (N12, N16 bars) with a ribbed profile. The tensile strength of these bars was nominally 500 MPa, which equates to a failure load of 56.5 kN and 100.5 kN respectively. The links comprised U6 mm stainless steel smooth bar with a tensile strength of 830 MPa.
2.2. Materials The mix design is shown in Table 1. The cement conformed to AS 3972 type GP and the aggregate basalt was of a specific gravity 2.99. The coarse and fine aggregate were prepared in accordance with AS 1141-2000. Salt (NaCl) was added to the mix at a rate of 3% by weight of cement. Mixing was undertaken in accordance with AS 1012.2-1994. Specimens were cured for 28 days under wet hessian before testing.
2. Materials and methods 2.1. Specimens Beam end specimens were selected for this study [2,24,25]. This type of eccentric pullout or ‘beam end’ type specimen uses a bonded length representative of the anchorage zone of a typical simply supported beam [26,27]. This configuration as opposed to the direct RILEM pullout test, which has appreciably higher cover than the minimum values prescribed in design codes, has been adopted in order to investigate the failure of the bar-concrete interface in a region where there are cracks due to corrosion and tensile and shear stress distributions. In an RC beam these cracks are radial around (and close to) the bar and extend to the surface as cracks both transverse and parallel to the span of the beam. The pullout failure is dependent upon the presence of these cracks, the confinement due to reinforcement stirrups and concrete cover and the stress distribution (flexural and shear stresses). The specimens adopted are designed to explore these interactions. Whilst it is accepted that the bond stress varies along the length of the
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Fig. 1. Beam end specimen design.
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Table 1 Concrete mix design. Material Quantity
Cement
w/c 3
381 kg/m
0.49
Sand
10 mm washed aggregate 3
517 kg/m
Restraints
3
463 kg/m
7 mm washed aggregate 463 kg/m
3
Salt 18.84 kg/m
Slump 3
140 ± 25 mm
Coupling Mechanism LVDT
MTS Acuator
Fig. 2. Bond Pull Out Test.
In order to compare bond strength for the different concrete compressive strengths, Eq. (1) is used to normalize bond strength for non-corroded specimens as has been used by other researchers [28].
s
0
sffiffiffiffiffiffi 40 ¼ sexp tl fc
ð2Þ
where s0 is the bond strength for grade 40 concrete, sexp tl is the experimental bond strength and f c is the experimental compressive strength. 2.3. Experimental methodology The time to onset of corrosion and the subsequent propagation of corrosion up to the appearance of external signs of structural distress normally takes many years. Consequently, accelerated corrosion has been used by many researchers to replicate the natural corrosion of reinforcing steel in concrete. These have involved experiments using impressed currents or artificial weathering with wet/dry cycles and elevated temperatures [1,8,10,13,14,18,29,30]. Corrosion rates (currents) measured in the field on actual structures are generally less than 5 lA/cm2 [31]. The maximum ‘natural’ current density reported from laboratory studies is approximately 100–200 lA/cm2 [32], though the bulk of the data reported in the literature are less than 10 lA/cm2. The majority of previous studies have used such a technique with current densities that are 3–100 times greater than the maximum value of 100 lA/cm2 observed under ‘natural’ conditions [33]. However, the current density levels are known to affect the surface crack width, because low current density levels give corrosion products the time to spread through the concrete pores and relieve the stress [6]. However, research [33] suggests that current densities up to 200 lA/cm2 result in similar stresses to 100 lA/cm2 during the early stages of corrosion. Clarke and Saifullah have suggested that current densities less than 250 lA/cm2 should be employed for accelerated tests as higher densities may have an adverse effect of the steel-concrete bond
[34], while Andrade et al. have used previously reported data with a maximum current density of 200 lA/cm2 to develop a model to estimate corrosion levels from crack width [32]. As such the work reported here has adopted a current density of 100 lA/cm2. While this is above the 5 lA/cm2 reported from field studies it does fall within the range adopted in previous accelerated studies corresponding the maximum value of 100 lA/cm2 observed under ‘natural’ conditions. It should be noted that the acceleration process however that the rate of corrosion may impact on the corrosion products, such that different oxidation state products may be formed, which could influence the bond. In the impressed current arrangement the steel bars served as the anode and four mild steel metal plates were fixed on the surface to serve as cathodes. Sponges (sprayed with salt water) were placed between the metal plates and the concrete surface to provide an adequate contact. The specimens were located within a custom designed accelerating system tank with a wet dry cycle of two one-hour wet cycles per 24 h. The first specimen was placed in the tank with the duration of impressed current recorded on each bar until the appearance of the first visible crack. A crack width value of 0.05 mm was considered to be the first visible crack. The average time was taken for the top two bars and the bottom two bars. This provided a reference time for determining preand post-cracking levels of corrosion as had been adopted in previous research [35]. This duration was then reduced to 1/3 and 2/3 of the average value for the remaining specimens to ensure they had undergone the early stages of corrosion prior to the appearance of any visible cracks – the main focus of this study. The current was disconnected from each bar in turn as the required (target) corrosion time was achieved. Consequently samples were tested at four stages – with no corrosion, at first visible crack, at one third and two thirds of the time to first visible crack (with different times for top and bottom bars). After the period of accelerated corrosion bond strength tests were conducted by means of a custom-built horizontally-acting test rig incorporating an MTS numerically controlled actuator. As noted above, plastic tube of length 80 mm was provided at the
Table 2 16 mm bars – corrosion loss (%) and bond strength (MPa). Position
None None None None 1/3 1/3 1/3 1/3 2/3 2/3 2/3 2/3 Crack Crack Crack Crack
16–3-c
16–3-u
16–2-c
Current Applied (Hours)
Actual % Loss
Bond (MPa)
Current Applied (Hours)
Actual % Loss
Bond (MPa)
Current Applied (Hours)
0 0 0 0 46 46 17.5 17.5 92 92 35 35 135 145 58 47
0.25 0.08 0.23 0.17 0.49 0.42 0.19 0.06 0.78
7.86
0 0 0 0 18 18 13 13 36 36 26.5 26.5 52 56 40 40
0.44 0.15 0.15 0.27 0.40 0.57 0.59 0.27 0.49 0.19 0.53 0.59 0.21 0.27 0.27 0.32
7.49
0 0 0 0 43 43 15 15 86 86 30 30 121 137 90 94
0.23 0.27 1.08 1.27 0.72 0.93
2.82 3.06 3.49 5.92
7.01 6.17
3.24 6.41
7.23 7.40 6.41
4.91 4.69 6.38 5.73
16–2-u Actual % Loss
Bond (MPa)
1.03
2.38
1.37 1.37
4.24 3.47
1.79
5.01
2.34
3.78
1.98 2.32
5.26 3.52
2.30
5.33
Current Applied (Hours) 0 0 0 0 81 81 66 66 162 162 132 132 205 281 181 215
16–1-c Actual % Loss
Current Applied (Hours)
Bond (MPa)
Current Applied (Hours)
3.11
0 0 0 0 32 32 43.5 43.5 64 64 87 87 102 90 140 120
2.79 3.02
0 0 0 0 108 108 64 64 216 216 127 127 349 299 184 198
2.04 1.56 1.52
2.74 2.68
1.92 1.79
3.60 3.57
2.14
1.60
16–1-u
Bond (MPa)
3.44
2.43 3.01
2.72 3.02
3.28 3.00
Bond (MPa)
4.90 4.82 5.69 5.66
5.49 5.47
4.95 5.31
Note: * 16–3-u corresponds to 16 mm diameter bar, 3 times cover, unconfined.
Table 3 12 mm bars – corrosion loss (%) and bond strength (MPa). 12–3-u
12–3-c
12–2-u
12–2-c
Position
Current Applied (Hours)
Current Applied (Hours)
Actual % Loss
Bond (MPa)
Current Applied (Hours)
Actual % Loss
Bond (MPa)
Current Applied (Hours)
Actual % Loss
Top Top Bottom Bottom Top Top Bottom Bottom Top Top Bottom Bottom Top Top Bottom Bottom
None None None None 1/3 1/3 1/3 1/3 2/3 2/3 2/3 2/3 Crack Crack Crack Crack
0 0 0 0 30.5 30.5 40 40 61 61 80 80 86 98 114 126
0.11 0.41 0.07 0.04 0.26 0.68 0.53 0.38 0.83 0.75 0.56 0.49 0.90 0.83 0.98 1.16
6.20
0 0 0 0 32.5 32.5 32.5 32.5 65 65 65 65 125 135 150 110
0.19 0.08 0.11 0.26 0.94 0.64 0.41 0.49 0.68 1.58 0.34 0.60 0.45 0.79 2.82 2.48
5.78 6.00 5.53
0 0 0 0 62 62 68 68 124 124 136 136 180 192 200 208
0.00 0.00 0.04 0.04 0.26 0.60 0.26 0.90 0.52 1.16 0.52 1.76
5.60 5.95 6.27 6.27 5.84 6.23 6.23 6.04 5.61 6.11
5.67 5.07 4.81 5.11 5.65 5.81 6.54 5.46 4.93 4.78 5.93
Bond (MPa)
2.75 5.77 3.38 4.44 2.40 3.25 1.24 3.69
1.54
4.59
12–1-c
12–1-u
Current Applied (Hours)
Actual % Loss
Bond (MPa)
Current Applied (Hours)
Bond (MPa)
Current Applied (Hours)
Bond (MPa)
0 0 0 0 62.5 62.5 80 80 125 125 160 160 190 184 220 260
0.00 0.00 0.04 0.04 0.56 1.09 0.75 0.60 1.91 0.41 1.12 1.16 1.05 0.86 1.65 1.76
3.33
0 0 0 0 18 18 22.5 22.5 36 36 45 45 54 54 80 75
2.84
0 0 0 0 67.5 67.5 52 52 135 135 104 104 195 210 165 147
3.64
4.94 3.91 5.58
3.82 4.21 4.19
5.26
3.38
2.87 2.59
2.59 2.28 2.71 2.43
3.68
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Top Top Bottom Bottom Top Top Bottom Bottom Top Top Bottom Bottom Top Top Bottom Bottom
Exposure
3.83 3.91 3.88 2.29
4.04 3.76
Note: * 12–3-u corresponds to 12 mm diameter bar, 3 times cover, unconfined. 553
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end of the concrete section underneath the transverse reaction to ensure that the bond strength was not enhanced by the reactive (compressive) force (acting normal to the bar). The specimen was positioned so that an axial force and no bending moment was applied to the bar being tested. The restraints were sufficiently rigid to ensure minimal rotation or twisting of the specimen during loading, Fig. 2. A uniform loading rate of 10kN/min was applied to the specimens. Top bars were tested first then the specimen was inverted to allow testing of the bottom bars. The bars were initially cleaned prior to casting with a 12% nitric acid solution, then washed in distilled water and neutralized by a calcium hydroxide solution before being washed in distilled in accordance with standard practice to ensure that all residue was removed [35]. Following the pull-out tests the corroded bars were cleaned in the same way and weighed again. The average observed corrosion penetration was calculated using the pre and post test weights and the theoretical corrosion by calculation using Faraday’s constant. Corrosion was assumed to occur only over the length of embedded bar exposed to the concrete (300 mm). The corrosion degree was determined using the following equation
CR ¼
ðG0 GÞ 100% g0 l
ð3Þ
where G0 is the initial weight of the steel bar before corrosion, G is the final weight of the steel bar after removal of the post-test corrosion products, g0 is the weight per unit length of the steel bar (0.888 and 1.58 g/mm for U12 mm and U16 mm bars respectively), l is the embedded bond length. 3. Results 3.1. Visual inspection The observed failure modes from the pull-out tests were splitting failure and splitting induced pullout failure. In the cases of the unconfined specimens the low cover resulted in the tension forces induced by the deformed bars surpassing the capacity of the surrounding concrete, resulting in the cover being sheared off from the surface, otherwise known as splitting failure. In the
confined cases, visible splitting cracks appeared coupled with an increase in bar slip, typically known as splitting induced pull-out failure. The failure mode (and associated failure load) is representative of that which would be expected in actual beams with corrosion of reinforcement present where insufficient reinforcement bond results in failure. As discussed above (Section 2.1) the failure load, expressed as a mean bond strength, is not comparable to that obtained from a direct pullout test rather it offers a means of assessing the loss of performance as a result of the most likely failure mechanism. It was noted that for confined specimens, the shear reinforcement showed significant signs of corrosion and pitting. This demonstrates that the confinement attracted impressed current away from the main reinforcing bars. 3.2. Corrosion loss Tables 2 and 3 show the duration of the applied current (hours) and the actual measured mass loss (%), while Fig. 3 shows the measured loss vs the theoretical loss. The theoretical loss was evaluated from the corrosion current, the time and Faraday’s law. Bars that were damaged in the extraction and cleaning process have been omitted. The mass loss shows no significant variation with cover, this is contrary to expectation where it would be anticipated that due to the confinement provided by the additional cover greater stress would be required to initiate cracking. Hence, a greater volume of rust product and a greater mass loss. A greater mass loss is however, required for confined specimens compared to unconfined specimens. This is attributed to the confining links resulting in a greater stress being required to induce cracking. Thus a greater quantity of rust is required to provide the increased stress, and hence the greater mass loss required to produce the greater volume of corrosion products. It was also noted that the total current passed to achieve cracking does not always correspond to the trends observed in the actual mass loss. In particular a relatively small quantity of current was required to initiate cracking in the 12-u1 and 12-c-1 specimens compared to the 2 and 3 times cover specimens, while in other cases similar durations were needed for specimens of different cover. In addition it was noted that actual loss for confined bars was 43% less than theoretical predictions and 33% less for the unconstrained bars.
Fig. 3. Mass loss (%), Theoretical and Actual.
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(a) 12mm and 16mm bars effect of confinement and bar posion
(b) Effect of confinement
(c) Effect of bar size
Fig. 4. Bond strength versus cover.
Fig. 5. Actual Mass Loss (%) vs Bond Strength, 16 mm bars, 3 times cover. Fig. 6. Actual Mass Loss (%) vs Bond Strength, 16 mm bars, 2 times cover.
These variations in the theoretical data are attributed to stray currents – i.e. current that is flowing through paths that circumvent the bar. This includes current that results in corrosion of the reinforcing links (where present). Despite being stainless steel there were signs of corrosion of these links, thus explaining the proportionally greater additional loss compared to the theoretical for the confined bars. The observation that the discrepancy is greater in theory as opposed to practice supports the idea that the constrained specimens suffered additional ‘loss of current’ that could be associated with corrosion of the links. The mean measured losses differed between the top bars and the bottom bars
with the bottom bars exhibiting 12% more in the confined specimens and 28% more in the unconfined specimens. This greater difference in the unconfined bars can be explained by the top unconfined bars exhibiting relatively poorer compaction and hence poorer containment by the surrounding concrete. This relative difference between top and bottom bars would be expected to reduce when confinement by stirrups is included. The smaller difference in values between top and bottom in the cases of the constrained specimens could also be evidence of the effect of the electrical connectivity between the top and bottom as a result of the links. This
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Fig. 7. Theoretical Mass Loss (%) vs Bond Strength, 16 mm bars, 1 times cover. Fig. 10. Theoretical Mass Loss (%) vs Bond Strength, 12 mm bars, 1 times cover.
considering the method of initial assessment of the time to achieve first crack. Each bar was disconnected when it reached first crack but the others were continued. In the case of confined specimens the links result in an electrically connected reinforcement cage and hence the disconnected bars may still be receiving current from the connected bars. Hence, explaining the reduced level of variation in corrosion loss between top and bottom bars in the constrained specimens. 3.3. Bond strength
Fig. 8. Actual Mass Loss (%) vs Bond Strength, 12 mm bars, 3 times cover.
The peak bond strength was determined by dividing the maximum load by the embedded area of the bar. This calculation was based on the assumption that the bond stress was uniform along the entire embedment length of bar encased within the concrete specimen. Results in Tables 2 and 3 identify that not all bars were tested which is as a result of damage incurred to the concrete specimens during previous pull-out tests from the same specimens. If the specimen base surface is damaged due to the destructive failure of the top face of the specimens in the first tests, the specimen cannot be adequately supported to test the third and/or fourth bars. Likewise if a bar was damaged in extraction the post test corrosion loss could not be assessed. 4. Discussion 4.1. Effect of casting position
Fig. 9. Actual Mass Loss (%) vs Bond Strength, 12 mm bars, 2 times cover.
further supports the premise that the stirrups (that link the top and bottom bars physically and electrically) may be playing a part in the forced corrosion process. This can be further explained by
Considering all bars (unaccelerated and accelerated) the bond strength of the bottom cast bars was generally greater than that of the top cast bars. In the case of the unaccelerated control specimens, where a small degree of corrosion was observed due to the NaCl within the mix a mean increase in bond strength of 14.5% was observed, similar to that reported for other authors for none corroded bars [21,34,36,37]. In the accelerated specimens the increase in the mean strength was 9% for the 16 mm bars and 10% for the 12 mm bars. These observations are in agreement with other authors [34,36–38]. It is generally accepted that corroded bottom cast bars have significantly improved bond compared to top cast bars due to the corrosion products filling the voids that are often present under top cast bars as the corrosion progresses [24]. The corrosion also acts as an ‘anchor’, similar to the ribs on deformed bars, to increase the bond. As cover increases it would be anticipated that greater compaction
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would be achieved around the top cast bars. Thus the area of voids would be reduced and thus the effect of the corrosion product filling these voids and increasing the bond strength would be reduced. Fig. 4 shows the influence of cover. There is a general increase in bond strength with cover as expected with the increase being more rapid in the case of unconfined as opposed to confined specimens. Fig. 4(a) demonstrates the range of values at the three different covers and shows that the variation between the top and bottom cast bars is reduced as cover increases, thus supporting the hypothesis above regarding the voids beneath the top cast bars. 4.2. Effect of bar size and confinement Fig. 4(c) demonstrates that there is little scaling effect with regard to cover between the 12 and 16 mm bars as the mean values of bond strength do not vary significantly over the three values of cover. The 16 mm bars exhibit a slightly higher value at one times cover, a slightly lower at two times cover and are effectively equal at three times cover. The mean values of bond strength for the unconfined and the confined cases are shown in Fig. 4(b). The variations show that as cover is increased the unconfined values of bond increase more rapidly than the confined values. This is to be expected as the influence of the confining stainless steel links becomes less significant as there is greater confinement from the concrete cover. Chana noted an increase in bond strength due to stirrups of between 10 and 20%. The trends suggest that at 3 times cover the influence of this level of confining reinforcement is minimal [24]. 4.3. Effect of degree of corrosion The variation in bond strength with degree of corrosion is shown in Figs. 5–10 for both 12 mm and 16 mm bars with covers of 1, 2 and 3 times bar diameter. The actual measured mass loss is shown, Figs. 5, 6, 8 and 9, corresponding to the 12 mm and 16 mm bars with 3 and 2 times cover, in the case of the 1 times cover the bars were damaged in the testing and the theoretical mass loss is shown to illustrate the trends, which are consistent for both the measured and theoretical mass loss. The confined specimens show a general increase for both the 16 mm at 1 and 2 times cover, while the 3 times cover show no clear trend. This would correlate with previous research which
557
has shown an increase in bond strength for confined specimens in the early stages of cracking [9,16,21,25,39]. The results would indicate that an increase in bond occurs once the corrosion products start to form. This is explained by considering the growth of cracking around the bar (even before visible cracks are present at the surface). In the initial stages of corrosion virtually all the dissolved iron ions react to form expansive corrosion products and an increase in compressive stress (pressure) around the bar. This results in an increase in the confinement and mechanical interlocking around the bar, and an increased roughness of the bar (due to the corrosion products) attributed to increased radial stresses on the bar–concrete interface in the vicinity of the end reactions [1,12,40]. However the growth of corrosion products around the bars also results in tensile hoop stresses and cracking that makes it possible for the iron ions to be transported along the crack and ultimately, as the cracks reach the surface, out of the concrete, thus reducing the bond strength. At higher cover the effect of the increased confinement due to this increased cover means that the corrosion products provide little if any increase in compressive stress around the bar. For the unconfined specimens the results are less clear cut with an initial increase noted in the 1 and 2 times cover, 16 mm bars, similar to that for the confined specimens, while for the 1 and 2 times cover 12 mm bars a decrease in bond strength is observed from the outset. In the 3 times cover an initial increase is observed but at higher mass loss a subsequent decrease is noted. The decrease noted is consistent with previous data where a decrease in bond strength from the onset is cracking is observed for unconfined bars. There is no clear explanation as to why the 16 mm and 12 mm bars display contrasting behavior at the 1 and 2 times cover. A possible explanation may be that the ribbed profile is greater in 16 mm bars and at very early stages the corrosion products may act as increased anchorage which gives an increase in bond strength, but this does not occur in the 12 mm bars. At 3 times cover the results can again be explained by the greater cover meaning that the corrosion product provides no increase in the compressive stress around the bar. Previous work using the same experimental arrangement focused on the variation in bond strength after visible surface cracking [25,39]. Fig. 11 shows the mean data from both trials. It can be seen that the results of this study correlate well with these results. In the 1 times cover specimens there is a small increase in bond strength before visible cracking, the strength then reduces
Unconfined trend
Confined trend
Fig. 11. Actual loss, 16 mm bars, 3 times cover, incl. previous post cracking results (Tang D et al., 2008).
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significantly as corrosion proceeds and surface cracks grow. In the 3 times cover specimens a reduction in strength from the first crack is observed. These results indicate that in the unconfined specimens the reduction in strength actually begins prior to cracking. Figs. 4–10 also demonstrate that at low cover the influence of the steel confinement (links) is more significant than at higher covers. This is also demonstrated in Fig. 4(b) and as discussed above is explained by the effect of the links becoming less significant as the confining contribution of the concrete cover becomes more significant. Other researchers have reported enhancements of bond stress of between 40% and 60% due to confinement. [2,18,41]. In uncorroded specimens Chana noted an increase in bond strength due to stirrups of between 10 and 20%. However the loading techniques have not all been the same and the results were based on the post-cracking phase of corrosion. Rodriguez et al. (1994) undertook pullout tests on confined and unconfined samples on similar beam end specimens as used here. The tests included cases with the same size bars as used above (16 mm) although with a smaller embedded length (210 mm as opposed to 300 mm). Based on the results an empirical relationship between bond strength, degree of corrosion, bar size, cover, link details and tensile strength of the concrete was proposed. The results show that for the links adopted here there would be an enhancement of bond strength due to confinement of approximately 2.0 MPa. This level of increase in bond strength is reflected in Figs. 4b, 5 and 8 at low concrete cover.
5. Conclusions 1. The data indicates that there is a relationship between extent of corrosion (measured by mass loss) and bond strength during the early stages of corrosion, before visible cracks occur. For confined specimens at one and two times cover, bond strength initially increases as corrosion proceeds. At three times cover, bond strength remains constant when corrosion commences due to the increased cover providing no increase in the compressive stress around the bar. 2. In unconfined specimens there is no clear relationship at one and two times cover with 12 mm bars displaying a reduction in bond strength from the commencement of corrosion, while 16 mm bars show an initial increase in bond strength prior to a decrease as the level of corrosion increases prior to cracking. At three times cover both 12 and 16 mm bars display this initial increase followed by a decrease as corrosion levels become greater prior to cracking. 3. The bond strength of the bottom cast bars was generally greater than that of the top cast bars. 4. The variation in bond strength between the top and bottom cast bars is reduced as cover increases, thus supporting the hypothesis that the extent of voids under the top cast bars is reduced as cover is increased. 5. The influence of shear links on bond strength is to increase confinement resulting in greater bond strength, this effect reduces as cover increases. 6. The effect of confinement (shear links) on the pull-out strength of bars with a cover of 3 times diameter is not significant. This may be due to the presence of the increased significance of the confinement resulting from the greater concrete cover. 7. Crack patterns both due to corrosion of specimens and during pull-out testing can differ significantly for specimens with/ without shear links. 8. There is little scaling effect in terms of bond strength with regard to cover between the 12 and 16 mm bars. 9. The shear links connecting the top and bottom bars physically and electrically played a significant role in the forced corrosion process resulting in corrosion of the links and a reduced level of
variation in corrosion loss between top and bottom bars in the constrained specimens.
6. References [1] J. Cairns et al., Mechanical properties of corrosion-damaged reinforcement, ACI Mater. J. 102 (4) (2005) 256–264. [2] J. Rodriguez, et al., Corrosion of reinforcement and service life of concrete structures, in Duability of Building materials and Components. 1996, E &FN Spon: London. p. 117-126. [3] J. Cairns, Bond and anchorage of embedded steel reinforcement in fib Model Code 2010, Struct. Concrete 16 (1) (2015) 45–55. [4] M. Prieto, P. Tanner, C. Andrade, Multiple linear regression model for the assessment of bond strength in corroded and non-corroded steel bars in structural concrete, Mater. Struct. 49 (2016) 4749–4763. [5] C. Andrade, C. Alonso, F.J. Molina, Cover cracking as a function of bar corrosion. 1. Experimental test, Mater. Struct. 26 (162) (1993) 453–464. [6] C. Andrade, et al., Cover cracking and amount of rebar corrosion; importance of the current applied in accelerated tests, in 4th International Congress Concrete in the Service of Mankind;Concrete Repair, Rehabilitation and Protection. 1996: Dundee. [7] C. Arya, F.K. OforiDarko, Influence of crack frequency on reinforcement corrosion in concrete, Cem. Concr. Res. 26 (3) (1996) 345–353. [8] C. Alonso et al., Factors controlling cracking of concrete affected by reinforcement corrosion, Mater. Struct. 31 (211) (1998) 435–441. [9] T. Vidal, A. Castel, R. Francois, Analyzing crack width to predict corrosion in reinforced concrete, Cem. Concr. Res. 34 (1) (2004) 165–174. [10] K. Vu, M.G. Stewart, J. Mullard, Corrosion-induced cracking: experimental data and predictive models, ACI Struct. J. 102 (5) (2005) 719–726. [11] F.J. Molina, C. Alonso, C. Andrade, Cover cracking as a function of rebar corrosion. 2. Numerical-model, Mater. Struct. 26 (163) (1993) 532–548. [12] J. Cairns, Y. Du, D. Law, Influence of corrosion on the friction characteristics of the steel/concrete interface, Constr. Build. Mater. 21 (1) (2007) 190–197. [13] T. El Maaddawy, K. Soudki, T. Topper, Long-term performance of corrosiondamaged reinforced concrete beams, ACI Struct. J. 102 (5) (2005) 649–656. [14] C. Fischer, J. Ozbolt, An appropriate indicator for bond strength degradation due to reinforcement corrosion, in VIII International Conference on Fracture Mechanics of Concrete and Concrete Structures, J.G.M. Van Mier, et al., Editors. 2012: Toeldo, Spain. pp. 1–5. [15] E.T. Maddawy, A.S. Khaled, Effectiveness of impressed current technique to simulate corrosion of steel reinforcement in concrete, J. Mater. Civ. Eng. 15 (1) (2003) 41–47. [16] C. Fischer, J. Ozbolt, Influence of bar diameter and concrete cover on bond degradation due to corrosion, in Bond in Concrete 2012: 4th Bond in Concrete Conference. 2012: Brescia, Italy. pp. 445–451. [17] X.H. Wang, X.L. Liu, Modeling bond strength of corroded reinforcement without stirrups, Cem. Concr. Res. 34 (8) (2004) 1331–1339. [18] C.Q. Fang et al., Bond behaviour of corroded reinforcing steel bars in concrete, Cem. Concr. Res. 36 (10) (2006) 1931–1938. [19] H.S. Lee, T. Noguchi, F. Tomosawa, Evaluation of the bond properties between concrete and reinforcement as a function of the degree of reinforcement corrosion, Cem. Concr. Res. 32 (8) (2002) 1313–1318. [20] A.A. Almusallam et al., Effect of reinforcement corrosion on bond strength, Constr. Build. Mater. 10 (2) (1996) 123–129. [21] D.W. Law et al., Impact of crack width on bond: confined and unconfined rebar, Mater. Struct. 44 (7) (2011) 1287–1296. [22] F.M. Li, Y.S. Yuan, Effects of corrosion on bond behavior between steel strand and concrete, Constr. Build. Mater. 38 (2013) 413–422. [23] T.C.K. Molyneaux, D.W. Law, D. Tang, The relationship between bond strength and surface crack width in reinforced concrete subject to chloride induced corrosion, in European Bridge Conference, 15th International Conference, M. Forde (ed.). 2014: Edinburgh. p. Paper 1224. [24] P.S. Chana, A test method to establish realistic bond stresses, Mag. Concr. Res. 42 (151) (1990) 83–90. [25] D.L. Tang et al., Influence of surface crack width on bond strength of reinforced concrete, ACI Mater. J. 108 (1) (2011) 29–37. [26] J. Cairns, Y. Du, D. Law, Residual bond strength of corroded plain round bars, Mag. Concr. Res. 58 (4) (2006) 221–231. [27] J. Cairns, G.A. Plizzari, Towards a harmonised European bond test, Mater. Struct. 36 (262) (2003) 498–506. [28] M.I. Baldwin, L.A. Clark, The assessment of reinforcing bars with inadequate anchorage, Mag. Concr. Res. 47 (171) (1995) 95–102. [29] J.A. Gonzalez et al., Comparison of rates of general corrosion and maximum fitting penetration on concrete embedded steel reinforcement, Cem. Concr. Res. 25 (2) (1995) 257–264. [30] D. Coronelli, K.Z. Hanjari, K. Lundgren, Severely corroded RC with cover cracking, J. Struct. Eng.-ASCE 139 (2) (2013) 221–232. [31] D.W. Law, F. Blin, K. Wilson, Interpretation and Modelling of Corrosion Rate Measurements on Reinforced Concrete Structures, in 4th International Conference on Non Destructive Monitoring. 2007: Harrogate, UK. pp. 2276– 2287. [32] C. Andrade et al., Estimation of corrosion attack in reinforced concrete by means of crack opening, Struct. Concr. (2016).
D.W. Law, T.C.K. Molyneaux / Construction and Building Materials 155 (2017) 550–559 [33] A.E.T. Maaddawy, A.S. Khaled, Effectiveness of Impressed Current Technique to Simulate Corrosion of Steel Reinforecement in Concrete. J. Mater. Civ. Eng., 2003, 15(1). [34] L.A. Clark, M. Saifullah, Effect of corrosion on reinforcement bond strength, in 5th Structural Faults and Repair Conference, M. Forde (ed.), 1990: Edinburgh. pp. 113–119. [35] C.Q. Fang et al., Corrosion influence on bond in reinforced concrete, Cem. Concr. Res. 34 (11) (2004) 2159–2167. [36] G.J. Alsulaimani et al., Influence of corrosion and cracking on bond behavior and strength of reinforced-concrete members, ACI Struct. J. 87 (2) (1990) 220– 231. [37] P.S. Mangat, M.S. Elgarf, Bond characteristics of corroding reinforcement in concrete beams, Mater. Struct. 32 (216) (1999) 89–97.
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[38] J. Cairns, Z. Zhao, Behavior of concrete beams with exposed reinforcement, Proc. Institut. Civil Eng.-Struct. Build. 99 (2) (1993) 141–154. [39] D. Tang, et al., The influence of chloride induced corrosion cracks on the strength of confined and unconfined reinforced concrete, in 12th International Conference Structural Faults & Repair, M. Forde (ed.), 2008: Edinburgh. [40] J. Cairns, Y. Du, D. Law, Structural performance of corrosion-damaged concrete beams, Mag. Concr. Res. 60 (5) (2008) 359–370. [41] J. Rodriguez, L.M. Ortega, J. Casal, Corrosion of reinforcing bars and service life of reinforced concrete structutres, in Concrete Across Borders. 1994, Odense, Denmark. pp. 315–326.
Contents lists available at ScienceDirect
Construction and Building Materials journal homepage: www.elsevier.com/locate/conbuildmat
Impact of corrosion on bond in uncracked concrete with confined and unconfined rebar David W. Law ⇑, Thomas C.K. Molyneaux School of Civil Engineering, RMIT University, Melbourne, Australia
h i g h l i g h t s For confined specimens bond strength initially increases as corrosion proceeds. Unconfined specimens display no clear relationship. The variation between top and bottom cast bars reduced as cover increased. The effect of confinement reduced as cover increased. There was little scaling effect between the 12 and 16 mm bars.
a r t i c l e
i n f o
Article history: Received 30 January 2017 Received in revised form 16 August 2017 Accepted 18 August 2017
Keywords: Bond Corrosion Rebar Cover Concrete Cracking
a b s t r a c t This paper examines the variation in bond strength in corroding specimens, before surface cracks are visible. Reinforced concrete specimens containing 16 mm and 12 mm diameter bars at covers of 1, 2 and 3 times bar diameter, with and without confinement, were subjected to accelerated corrosion. Bond strength is seen to increase during the early stages of corrosion (pre external cracking) for confined bars at all covers and unconfined bars at 3 times cover. However, while 16 mm bars, 1 and 2 times cover showed similar behaviour the 12 mm bars displayed a decrease in bond once corrosion was initiated. Ó 2017 Elsevier Ltd. All rights reserved.
1. Introduction Corrosion of reinforcing steel in concrete structures is one of the primary causes of concern for owners with billions of pounds being spent on repair and maintenance. The bond between the steel reinforcement and the concrete is essential for the reinforced concrete to act in a composite manner and the presence of corrosion threatens this integrity. At present much of the inspection and remediation is conducted when there is visual evidence of corrosion, such as rust staining and cracking/spalling. However there is a significant period during which corrosion is continuing and the bond strength changing whilst there are no surface signs. Corrosion of reinforcing steel in concrete is a complex electrochemical process. Reinforcement in concrete is resistant to corrosion due to the protective iron oxide film that arises from
⇑ Corresponding author. E-mail address: [email protected] (D.W. Law). http://dx.doi.org/10.1016/j.conbuildmat.2017.08.112 0950-0618/Ó 2017 Elsevier Ltd. All rights reserved.
the alkaline conditions (a pH of approximately 12.5) provided by the surrounding concrete. Chloride ions however, from an external marine environment or from de-icing salt can impair the protective oxide film and initiate corrosion. Variations in properties at the surface of the bar encourage a mixed electrode to form in which one region of the steel acts as an anode and another as the cathode with pore water and dissolved ions forming an electrolytic cell. At the anode iron atoms pass into solution as ferrous ions (positively charged). For corrosion to occur a loss of passivation, the presence of oxygen and the presence of water are all required. The deterioration of reinforced concrete is characterized by a general or localized loss of section of the reinforcing bars and the formation of expansive corrosion products at the surface of the bar. The production of expansive products creates tensile hoop stresses around the bar within the concrete, which can result in cracking and spalling of the concrete cover. This cracking can lead to accelerated ingress of aggressive agents thus accelerating further corrosion. This process leads to loss of strength and stiffness
D.W. Law, T.C.K. Molyneaux / Construction and Building Materials 155 (2017) 550–559
of the concrete cover and threatens the bond strength between the concrete and the reinforcing steel. Finally the corrosion reduces the cross section of the reinforcing steel, which can affect the ductility of the steel and the load bearing capacity, which can ultimately impact upon the serviceability of the structure and the structural capacity [1,2]. During the corrosion process the volume of the corrosion products created on the surface of the bar is greater than that of the steel consumed during the process of corrosion and hence radial and circumferential stresses are generated in the concrete around the steel bars, resulting in the cracking of the surrounding concrete. The mass of steel lost due to a corrosion current (I) can be calculated by applying Faraday’s law.
Dm ¼ MIt=zF
ð1Þ
where Dm is the mass of steel consumed (g), M is the atomic weight of the metal (56 g for iron); t is the time in seconds; z is the ionic charge (2 for iron); and F is Faraday’s constant (96,500 A/s). There has been a significant effort in assessing and modeling the degradation of reinforced concrete structures in order to assist asset managers in adequately planning for the prevention/minimisation of corrosion damage and maintenance, with a number of models proposed to predict the effect of corrosion on bond [2–4]. Research has been reported on the relationship between the corrosion level (corrosion penetration or mass loss of the steel) and cracks [5–14]; current density versus surface crack width [1,15,16]; and bond strength versus corrosion level (mass loss or corrosion penetration). The use of shear reinforcement within the concrete can counteract the loss of bond capacity to a certain degree by providing confinement and resisting the radial tensile cracking. Work by Wang et al. (2004) shows that bond strength increases because of confining action given by stirrups [17–19]. However, relatively little research has studied the relationship between bond strength and surface crack width [16,20–22] with even less on the effect of corrosion on bond prior to cracking [23]. Previous research has shown an initial increase in bond strength for corroded confined when cracking first occurs, however, for unconfined bars a decrease in bond strength is observed with the initial cracking [16,21]. However the influence of the level of corrosion prior to cracking is not clear. This paper reports the effect of corrosion level on confined and unconfined bars of 12 mm and 16 mm diameter, with 1, 2 and 3 times cover up to the initiation of the first crack.
embedded bar, the mean stress at peak pullout force is representative of the relative structural capacity. Specimens of rectangular cross section were cast with a longitudinal reinforcing bar in each corner. Specimens with confined and unconfined 12 mm and 16 mm diameter bars were cast with covers of 1, 2 and 3 times bar diameter. The specimens were dimensioned: 380 mm length, 200 mm width and 300 mm depth with an 80 mm plastic tube placed on the far (unloaded) end of each steel reinforcing bar to create a de-bonded zone which protects the reinforcement from the confining pressure of the concrete due to the transverse reaction (supporting) force during the pullout testing, Fig. 1. Confinement, where applicable, comprised undeformed stainless steel 6 mm diameter links at 100 mm spacing, three links in each specimen, positioned symmetrically between the end of the plastic tube and the end of the specimen. The specimens are similar in form to those adopted by [24]. Four specimens of each type were cast making a total of 48 specimens (3 covers, 2 bar sizes, confined and unconfined). The longitudinal reinforcement selected comprised standard deformed bars of 12 mm and 16 mm diameter (N12, N16 bars) with a ribbed profile. The tensile strength of these bars was nominally 500 MPa, which equates to a failure load of 56.5 kN and 100.5 kN respectively. The links comprised U6 mm stainless steel smooth bar with a tensile strength of 830 MPa.
2.2. Materials The mix design is shown in Table 1. The cement conformed to AS 3972 type GP and the aggregate basalt was of a specific gravity 2.99. The coarse and fine aggregate were prepared in accordance with AS 1141-2000. Salt (NaCl) was added to the mix at a rate of 3% by weight of cement. Mixing was undertaken in accordance with AS 1012.2-1994. Specimens were cured for 28 days under wet hessian before testing.
2. Materials and methods 2.1. Specimens Beam end specimens were selected for this study [2,24,25]. This type of eccentric pullout or ‘beam end’ type specimen uses a bonded length representative of the anchorage zone of a typical simply supported beam [26,27]. This configuration as opposed to the direct RILEM pullout test, which has appreciably higher cover than the minimum values prescribed in design codes, has been adopted in order to investigate the failure of the bar-concrete interface in a region where there are cracks due to corrosion and tensile and shear stress distributions. In an RC beam these cracks are radial around (and close to) the bar and extend to the surface as cracks both transverse and parallel to the span of the beam. The pullout failure is dependent upon the presence of these cracks, the confinement due to reinforcement stirrups and concrete cover and the stress distribution (flexural and shear stresses). The specimens adopted are designed to explore these interactions. Whilst it is accepted that the bond stress varies along the length of the
551
Fig. 1. Beam end specimen design.
552
D.W. Law, T.C.K. Molyneaux / Construction and Building Materials 155 (2017) 550–559
Table 1 Concrete mix design. Material Quantity
Cement
w/c 3
381 kg/m
0.49
Sand
10 mm washed aggregate 3
517 kg/m
Restraints
3
463 kg/m
7 mm washed aggregate 463 kg/m
3
Salt 18.84 kg/m
Slump 3
140 ± 25 mm
Coupling Mechanism LVDT
MTS Acuator
Fig. 2. Bond Pull Out Test.
In order to compare bond strength for the different concrete compressive strengths, Eq. (1) is used to normalize bond strength for non-corroded specimens as has been used by other researchers [28].
s
0
sffiffiffiffiffiffi 40 ¼ sexp tl fc
ð2Þ
where s0 is the bond strength for grade 40 concrete, sexp tl is the experimental bond strength and f c is the experimental compressive strength. 2.3. Experimental methodology The time to onset of corrosion and the subsequent propagation of corrosion up to the appearance of external signs of structural distress normally takes many years. Consequently, accelerated corrosion has been used by many researchers to replicate the natural corrosion of reinforcing steel in concrete. These have involved experiments using impressed currents or artificial weathering with wet/dry cycles and elevated temperatures [1,8,10,13,14,18,29,30]. Corrosion rates (currents) measured in the field on actual structures are generally less than 5 lA/cm2 [31]. The maximum ‘natural’ current density reported from laboratory studies is approximately 100–200 lA/cm2 [32], though the bulk of the data reported in the literature are less than 10 lA/cm2. The majority of previous studies have used such a technique with current densities that are 3–100 times greater than the maximum value of 100 lA/cm2 observed under ‘natural’ conditions [33]. However, the current density levels are known to affect the surface crack width, because low current density levels give corrosion products the time to spread through the concrete pores and relieve the stress [6]. However, research [33] suggests that current densities up to 200 lA/cm2 result in similar stresses to 100 lA/cm2 during the early stages of corrosion. Clarke and Saifullah have suggested that current densities less than 250 lA/cm2 should be employed for accelerated tests as higher densities may have an adverse effect of the steel-concrete bond
[34], while Andrade et al. have used previously reported data with a maximum current density of 200 lA/cm2 to develop a model to estimate corrosion levels from crack width [32]. As such the work reported here has adopted a current density of 100 lA/cm2. While this is above the 5 lA/cm2 reported from field studies it does fall within the range adopted in previous accelerated studies corresponding the maximum value of 100 lA/cm2 observed under ‘natural’ conditions. It should be noted that the acceleration process however that the rate of corrosion may impact on the corrosion products, such that different oxidation state products may be formed, which could influence the bond. In the impressed current arrangement the steel bars served as the anode and four mild steel metal plates were fixed on the surface to serve as cathodes. Sponges (sprayed with salt water) were placed between the metal plates and the concrete surface to provide an adequate contact. The specimens were located within a custom designed accelerating system tank with a wet dry cycle of two one-hour wet cycles per 24 h. The first specimen was placed in the tank with the duration of impressed current recorded on each bar until the appearance of the first visible crack. A crack width value of 0.05 mm was considered to be the first visible crack. The average time was taken for the top two bars and the bottom two bars. This provided a reference time for determining preand post-cracking levels of corrosion as had been adopted in previous research [35]. This duration was then reduced to 1/3 and 2/3 of the average value for the remaining specimens to ensure they had undergone the early stages of corrosion prior to the appearance of any visible cracks – the main focus of this study. The current was disconnected from each bar in turn as the required (target) corrosion time was achieved. Consequently samples were tested at four stages – with no corrosion, at first visible crack, at one third and two thirds of the time to first visible crack (with different times for top and bottom bars). After the period of accelerated corrosion bond strength tests were conducted by means of a custom-built horizontally-acting test rig incorporating an MTS numerically controlled actuator. As noted above, plastic tube of length 80 mm was provided at the
Table 2 16 mm bars – corrosion loss (%) and bond strength (MPa). Position
None None None None 1/3 1/3 1/3 1/3 2/3 2/3 2/3 2/3 Crack Crack Crack Crack
16–3-c
16–3-u
16–2-c
Current Applied (Hours)
Actual % Loss
Bond (MPa)
Current Applied (Hours)
Actual % Loss
Bond (MPa)
Current Applied (Hours)
0 0 0 0 46 46 17.5 17.5 92 92 35 35 135 145 58 47
0.25 0.08 0.23 0.17 0.49 0.42 0.19 0.06 0.78
7.86
0 0 0 0 18 18 13 13 36 36 26.5 26.5 52 56 40 40
0.44 0.15 0.15 0.27 0.40 0.57 0.59 0.27 0.49 0.19 0.53 0.59 0.21 0.27 0.27 0.32
7.49
0 0 0 0 43 43 15 15 86 86 30 30 121 137 90 94
0.23 0.27 1.08 1.27 0.72 0.93
2.82 3.06 3.49 5.92
7.01 6.17
3.24 6.41
7.23 7.40 6.41
4.91 4.69 6.38 5.73
16–2-u Actual % Loss
Bond (MPa)
1.03
2.38
1.37 1.37
4.24 3.47
1.79
5.01
2.34
3.78
1.98 2.32
5.26 3.52
2.30
5.33
Current Applied (Hours) 0 0 0 0 81 81 66 66 162 162 132 132 205 281 181 215
16–1-c Actual % Loss
Current Applied (Hours)
Bond (MPa)
Current Applied (Hours)
3.11
0 0 0 0 32 32 43.5 43.5 64 64 87 87 102 90 140 120
2.79 3.02
0 0 0 0 108 108 64 64 216 216 127 127 349 299 184 198
2.04 1.56 1.52
2.74 2.68
1.92 1.79
3.60 3.57
2.14
1.60
16–1-u
Bond (MPa)
3.44
2.43 3.01
2.72 3.02
3.28 3.00
Bond (MPa)
4.90 4.82 5.69 5.66
5.49 5.47
4.95 5.31
Note: * 16–3-u corresponds to 16 mm diameter bar, 3 times cover, unconfined.
Table 3 12 mm bars – corrosion loss (%) and bond strength (MPa). 12–3-u
12–3-c
12–2-u
12–2-c
Position
Current Applied (Hours)
Current Applied (Hours)
Actual % Loss
Bond (MPa)
Current Applied (Hours)
Actual % Loss
Bond (MPa)
Current Applied (Hours)
Actual % Loss
Top Top Bottom Bottom Top Top Bottom Bottom Top Top Bottom Bottom Top Top Bottom Bottom
None None None None 1/3 1/3 1/3 1/3 2/3 2/3 2/3 2/3 Crack Crack Crack Crack
0 0 0 0 30.5 30.5 40 40 61 61 80 80 86 98 114 126
0.11 0.41 0.07 0.04 0.26 0.68 0.53 0.38 0.83 0.75 0.56 0.49 0.90 0.83 0.98 1.16
6.20
0 0 0 0 32.5 32.5 32.5 32.5 65 65 65 65 125 135 150 110
0.19 0.08 0.11 0.26 0.94 0.64 0.41 0.49 0.68 1.58 0.34 0.60 0.45 0.79 2.82 2.48
5.78 6.00 5.53
0 0 0 0 62 62 68 68 124 124 136 136 180 192 200 208
0.00 0.00 0.04 0.04 0.26 0.60 0.26 0.90 0.52 1.16 0.52 1.76
5.60 5.95 6.27 6.27 5.84 6.23 6.23 6.04 5.61 6.11
5.67 5.07 4.81 5.11 5.65 5.81 6.54 5.46 4.93 4.78 5.93
Bond (MPa)
2.75 5.77 3.38 4.44 2.40 3.25 1.24 3.69
1.54
4.59
12–1-c
12–1-u
Current Applied (Hours)
Actual % Loss
Bond (MPa)
Current Applied (Hours)
Bond (MPa)
Current Applied (Hours)
Bond (MPa)
0 0 0 0 62.5 62.5 80 80 125 125 160 160 190 184 220 260
0.00 0.00 0.04 0.04 0.56 1.09 0.75 0.60 1.91 0.41 1.12 1.16 1.05 0.86 1.65 1.76
3.33
0 0 0 0 18 18 22.5 22.5 36 36 45 45 54 54 80 75
2.84
0 0 0 0 67.5 67.5 52 52 135 135 104 104 195 210 165 147
3.64
4.94 3.91 5.58
3.82 4.21 4.19
5.26
3.38
2.87 2.59
2.59 2.28 2.71 2.43
3.68
D.W. Law, T.C.K. Molyneaux / Construction and Building Materials 155 (2017) 550–559
Top Top Bottom Bottom Top Top Bottom Bottom Top Top Bottom Bottom Top Top Bottom Bottom
Exposure
3.83 3.91 3.88 2.29
4.04 3.76
Note: * 12–3-u corresponds to 12 mm diameter bar, 3 times cover, unconfined. 553
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end of the concrete section underneath the transverse reaction to ensure that the bond strength was not enhanced by the reactive (compressive) force (acting normal to the bar). The specimen was positioned so that an axial force and no bending moment was applied to the bar being tested. The restraints were sufficiently rigid to ensure minimal rotation or twisting of the specimen during loading, Fig. 2. A uniform loading rate of 10kN/min was applied to the specimens. Top bars were tested first then the specimen was inverted to allow testing of the bottom bars. The bars were initially cleaned prior to casting with a 12% nitric acid solution, then washed in distilled water and neutralized by a calcium hydroxide solution before being washed in distilled in accordance with standard practice to ensure that all residue was removed [35]. Following the pull-out tests the corroded bars were cleaned in the same way and weighed again. The average observed corrosion penetration was calculated using the pre and post test weights and the theoretical corrosion by calculation using Faraday’s constant. Corrosion was assumed to occur only over the length of embedded bar exposed to the concrete (300 mm). The corrosion degree was determined using the following equation
CR ¼
ðG0 GÞ 100% g0 l
ð3Þ
where G0 is the initial weight of the steel bar before corrosion, G is the final weight of the steel bar after removal of the post-test corrosion products, g0 is the weight per unit length of the steel bar (0.888 and 1.58 g/mm for U12 mm and U16 mm bars respectively), l is the embedded bond length. 3. Results 3.1. Visual inspection The observed failure modes from the pull-out tests were splitting failure and splitting induced pullout failure. In the cases of the unconfined specimens the low cover resulted in the tension forces induced by the deformed bars surpassing the capacity of the surrounding concrete, resulting in the cover being sheared off from the surface, otherwise known as splitting failure. In the
confined cases, visible splitting cracks appeared coupled with an increase in bar slip, typically known as splitting induced pull-out failure. The failure mode (and associated failure load) is representative of that which would be expected in actual beams with corrosion of reinforcement present where insufficient reinforcement bond results in failure. As discussed above (Section 2.1) the failure load, expressed as a mean bond strength, is not comparable to that obtained from a direct pullout test rather it offers a means of assessing the loss of performance as a result of the most likely failure mechanism. It was noted that for confined specimens, the shear reinforcement showed significant signs of corrosion and pitting. This demonstrates that the confinement attracted impressed current away from the main reinforcing bars. 3.2. Corrosion loss Tables 2 and 3 show the duration of the applied current (hours) and the actual measured mass loss (%), while Fig. 3 shows the measured loss vs the theoretical loss. The theoretical loss was evaluated from the corrosion current, the time and Faraday’s law. Bars that were damaged in the extraction and cleaning process have been omitted. The mass loss shows no significant variation with cover, this is contrary to expectation where it would be anticipated that due to the confinement provided by the additional cover greater stress would be required to initiate cracking. Hence, a greater volume of rust product and a greater mass loss. A greater mass loss is however, required for confined specimens compared to unconfined specimens. This is attributed to the confining links resulting in a greater stress being required to induce cracking. Thus a greater quantity of rust is required to provide the increased stress, and hence the greater mass loss required to produce the greater volume of corrosion products. It was also noted that the total current passed to achieve cracking does not always correspond to the trends observed in the actual mass loss. In particular a relatively small quantity of current was required to initiate cracking in the 12-u1 and 12-c-1 specimens compared to the 2 and 3 times cover specimens, while in other cases similar durations were needed for specimens of different cover. In addition it was noted that actual loss for confined bars was 43% less than theoretical predictions and 33% less for the unconstrained bars.
Fig. 3. Mass loss (%), Theoretical and Actual.
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(a) 12mm and 16mm bars effect of confinement and bar posion
(b) Effect of confinement
(c) Effect of bar size
Fig. 4. Bond strength versus cover.
Fig. 5. Actual Mass Loss (%) vs Bond Strength, 16 mm bars, 3 times cover. Fig. 6. Actual Mass Loss (%) vs Bond Strength, 16 mm bars, 2 times cover.
These variations in the theoretical data are attributed to stray currents – i.e. current that is flowing through paths that circumvent the bar. This includes current that results in corrosion of the reinforcing links (where present). Despite being stainless steel there were signs of corrosion of these links, thus explaining the proportionally greater additional loss compared to the theoretical for the confined bars. The observation that the discrepancy is greater in theory as opposed to practice supports the idea that the constrained specimens suffered additional ‘loss of current’ that could be associated with corrosion of the links. The mean measured losses differed between the top bars and the bottom bars
with the bottom bars exhibiting 12% more in the confined specimens and 28% more in the unconfined specimens. This greater difference in the unconfined bars can be explained by the top unconfined bars exhibiting relatively poorer compaction and hence poorer containment by the surrounding concrete. This relative difference between top and bottom bars would be expected to reduce when confinement by stirrups is included. The smaller difference in values between top and bottom in the cases of the constrained specimens could also be evidence of the effect of the electrical connectivity between the top and bottom as a result of the links. This
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Fig. 7. Theoretical Mass Loss (%) vs Bond Strength, 16 mm bars, 1 times cover. Fig. 10. Theoretical Mass Loss (%) vs Bond Strength, 12 mm bars, 1 times cover.
considering the method of initial assessment of the time to achieve first crack. Each bar was disconnected when it reached first crack but the others were continued. In the case of confined specimens the links result in an electrically connected reinforcement cage and hence the disconnected bars may still be receiving current from the connected bars. Hence, explaining the reduced level of variation in corrosion loss between top and bottom bars in the constrained specimens. 3.3. Bond strength
Fig. 8. Actual Mass Loss (%) vs Bond Strength, 12 mm bars, 3 times cover.
The peak bond strength was determined by dividing the maximum load by the embedded area of the bar. This calculation was based on the assumption that the bond stress was uniform along the entire embedment length of bar encased within the concrete specimen. Results in Tables 2 and 3 identify that not all bars were tested which is as a result of damage incurred to the concrete specimens during previous pull-out tests from the same specimens. If the specimen base surface is damaged due to the destructive failure of the top face of the specimens in the first tests, the specimen cannot be adequately supported to test the third and/or fourth bars. Likewise if a bar was damaged in extraction the post test corrosion loss could not be assessed. 4. Discussion 4.1. Effect of casting position
Fig. 9. Actual Mass Loss (%) vs Bond Strength, 12 mm bars, 2 times cover.
further supports the premise that the stirrups (that link the top and bottom bars physically and electrically) may be playing a part in the forced corrosion process. This can be further explained by
Considering all bars (unaccelerated and accelerated) the bond strength of the bottom cast bars was generally greater than that of the top cast bars. In the case of the unaccelerated control specimens, where a small degree of corrosion was observed due to the NaCl within the mix a mean increase in bond strength of 14.5% was observed, similar to that reported for other authors for none corroded bars [21,34,36,37]. In the accelerated specimens the increase in the mean strength was 9% for the 16 mm bars and 10% for the 12 mm bars. These observations are in agreement with other authors [34,36–38]. It is generally accepted that corroded bottom cast bars have significantly improved bond compared to top cast bars due to the corrosion products filling the voids that are often present under top cast bars as the corrosion progresses [24]. The corrosion also acts as an ‘anchor’, similar to the ribs on deformed bars, to increase the bond. As cover increases it would be anticipated that greater compaction
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would be achieved around the top cast bars. Thus the area of voids would be reduced and thus the effect of the corrosion product filling these voids and increasing the bond strength would be reduced. Fig. 4 shows the influence of cover. There is a general increase in bond strength with cover as expected with the increase being more rapid in the case of unconfined as opposed to confined specimens. Fig. 4(a) demonstrates the range of values at the three different covers and shows that the variation between the top and bottom cast bars is reduced as cover increases, thus supporting the hypothesis above regarding the voids beneath the top cast bars. 4.2. Effect of bar size and confinement Fig. 4(c) demonstrates that there is little scaling effect with regard to cover between the 12 and 16 mm bars as the mean values of bond strength do not vary significantly over the three values of cover. The 16 mm bars exhibit a slightly higher value at one times cover, a slightly lower at two times cover and are effectively equal at three times cover. The mean values of bond strength for the unconfined and the confined cases are shown in Fig. 4(b). The variations show that as cover is increased the unconfined values of bond increase more rapidly than the confined values. This is to be expected as the influence of the confining stainless steel links becomes less significant as there is greater confinement from the concrete cover. Chana noted an increase in bond strength due to stirrups of between 10 and 20%. The trends suggest that at 3 times cover the influence of this level of confining reinforcement is minimal [24]. 4.3. Effect of degree of corrosion The variation in bond strength with degree of corrosion is shown in Figs. 5–10 for both 12 mm and 16 mm bars with covers of 1, 2 and 3 times bar diameter. The actual measured mass loss is shown, Figs. 5, 6, 8 and 9, corresponding to the 12 mm and 16 mm bars with 3 and 2 times cover, in the case of the 1 times cover the bars were damaged in the testing and the theoretical mass loss is shown to illustrate the trends, which are consistent for both the measured and theoretical mass loss. The confined specimens show a general increase for both the 16 mm at 1 and 2 times cover, while the 3 times cover show no clear trend. This would correlate with previous research which
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has shown an increase in bond strength for confined specimens in the early stages of cracking [9,16,21,25,39]. The results would indicate that an increase in bond occurs once the corrosion products start to form. This is explained by considering the growth of cracking around the bar (even before visible cracks are present at the surface). In the initial stages of corrosion virtually all the dissolved iron ions react to form expansive corrosion products and an increase in compressive stress (pressure) around the bar. This results in an increase in the confinement and mechanical interlocking around the bar, and an increased roughness of the bar (due to the corrosion products) attributed to increased radial stresses on the bar–concrete interface in the vicinity of the end reactions [1,12,40]. However the growth of corrosion products around the bars also results in tensile hoop stresses and cracking that makes it possible for the iron ions to be transported along the crack and ultimately, as the cracks reach the surface, out of the concrete, thus reducing the bond strength. At higher cover the effect of the increased confinement due to this increased cover means that the corrosion products provide little if any increase in compressive stress around the bar. For the unconfined specimens the results are less clear cut with an initial increase noted in the 1 and 2 times cover, 16 mm bars, similar to that for the confined specimens, while for the 1 and 2 times cover 12 mm bars a decrease in bond strength is observed from the outset. In the 3 times cover an initial increase is observed but at higher mass loss a subsequent decrease is noted. The decrease noted is consistent with previous data where a decrease in bond strength from the onset is cracking is observed for unconfined bars. There is no clear explanation as to why the 16 mm and 12 mm bars display contrasting behavior at the 1 and 2 times cover. A possible explanation may be that the ribbed profile is greater in 16 mm bars and at very early stages the corrosion products may act as increased anchorage which gives an increase in bond strength, but this does not occur in the 12 mm bars. At 3 times cover the results can again be explained by the greater cover meaning that the corrosion product provides no increase in the compressive stress around the bar. Previous work using the same experimental arrangement focused on the variation in bond strength after visible surface cracking [25,39]. Fig. 11 shows the mean data from both trials. It can be seen that the results of this study correlate well with these results. In the 1 times cover specimens there is a small increase in bond strength before visible cracking, the strength then reduces
Unconfined trend
Confined trend
Fig. 11. Actual loss, 16 mm bars, 3 times cover, incl. previous post cracking results (Tang D et al., 2008).
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significantly as corrosion proceeds and surface cracks grow. In the 3 times cover specimens a reduction in strength from the first crack is observed. These results indicate that in the unconfined specimens the reduction in strength actually begins prior to cracking. Figs. 4–10 also demonstrate that at low cover the influence of the steel confinement (links) is more significant than at higher covers. This is also demonstrated in Fig. 4(b) and as discussed above is explained by the effect of the links becoming less significant as the confining contribution of the concrete cover becomes more significant. Other researchers have reported enhancements of bond stress of between 40% and 60% due to confinement. [2,18,41]. In uncorroded specimens Chana noted an increase in bond strength due to stirrups of between 10 and 20%. However the loading techniques have not all been the same and the results were based on the post-cracking phase of corrosion. Rodriguez et al. (1994) undertook pullout tests on confined and unconfined samples on similar beam end specimens as used here. The tests included cases with the same size bars as used above (16 mm) although with a smaller embedded length (210 mm as opposed to 300 mm). Based on the results an empirical relationship between bond strength, degree of corrosion, bar size, cover, link details and tensile strength of the concrete was proposed. The results show that for the links adopted here there would be an enhancement of bond strength due to confinement of approximately 2.0 MPa. This level of increase in bond strength is reflected in Figs. 4b, 5 and 8 at low concrete cover.
5. Conclusions 1. The data indicates that there is a relationship between extent of corrosion (measured by mass loss) and bond strength during the early stages of corrosion, before visible cracks occur. For confined specimens at one and two times cover, bond strength initially increases as corrosion proceeds. At three times cover, bond strength remains constant when corrosion commences due to the increased cover providing no increase in the compressive stress around the bar. 2. In unconfined specimens there is no clear relationship at one and two times cover with 12 mm bars displaying a reduction in bond strength from the commencement of corrosion, while 16 mm bars show an initial increase in bond strength prior to a decrease as the level of corrosion increases prior to cracking. At three times cover both 12 and 16 mm bars display this initial increase followed by a decrease as corrosion levels become greater prior to cracking. 3. The bond strength of the bottom cast bars was generally greater than that of the top cast bars. 4. The variation in bond strength between the top and bottom cast bars is reduced as cover increases, thus supporting the hypothesis that the extent of voids under the top cast bars is reduced as cover is increased. 5. The influence of shear links on bond strength is to increase confinement resulting in greater bond strength, this effect reduces as cover increases. 6. The effect of confinement (shear links) on the pull-out strength of bars with a cover of 3 times diameter is not significant. This may be due to the presence of the increased significance of the confinement resulting from the greater concrete cover. 7. Crack patterns both due to corrosion of specimens and during pull-out testing can differ significantly for specimens with/ without shear links. 8. There is little scaling effect in terms of bond strength with regard to cover between the 12 and 16 mm bars. 9. The shear links connecting the top and bottom bars physically and electrically played a significant role in the forced corrosion process resulting in corrosion of the links and a reduced level of
variation in corrosion loss between top and bottom bars in the constrained specimens.
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