Bond of corroded reinforcement in strain resilient cementitious composites

Sousana Tastani

doi:10.1515/corrrev-2022-0068

Article Open Access

Bond of corroded reinforcement in strain resilient cementitious composites

Sousana Tastani

Published/Copyright: February 9, 2023

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From the journal Corrosion Reviews Volume 41 Issue 2

Abstract

In corroded steel anchorages, radial pressures generated by corrosion and bond are balanced by hoop tension provided by the concrete cover. As the concrete has negligible tension strain capacity, in the absence of confinement, anchorage failure occurs by cover splitting and consequent bond deterioration. This paper presents an experimental study of short reinforcement anchorages embedded in a strain resilient cementitious composite matrix, which were preconditioned under simulated corrosion before mechanical loading. The experiment was meant to assess the restraining effect of randomly dispersed plastic fibers on the build-up of expansive rust, by measuring the degree of rebar corrosion that led to cracking of the concrete cover, and the consequent improvements in residual anchorage capacity, by measuring the mechanical response to four-point loading. A total of eighteen short beams were tested with variables being the anchorage length extending into the shear span of the beams as well as the matrix (with and without fibers). Comparisons among artificially corroded and non-corroded coupons demonstrate the favorable effect a strain resilient matrix has in delaying metal consumption as a consequence of the fibers’ restrain against crack initiation/widening; their mechanical testing reveals even increase of local bond strength for corrosion level less than 10%.

Keywords: bond; cementitious composites; steel reinforcement corrosion

1 Introduction

Corrosion of steel reinforcement, when it exceeds certain limits of mass loss, affects to varying degrees all response mechanisms, namely flexure, shear and anchorage/splice of a reinforced concrete (r.c.) member. The predominant failure mechanism arises from comparison between the competing strengths and their accompanying deformations; these indices are impaired by the bar area loss, the attenuation of rib profile, the steel embrittlement and the concrete cover cracking (Coronelli et al. 2019; El-Joukhadar et al. 2019; Goksu and Ilki 2016; Tastani and Pantazopoulou 2005).

In the case of corroded anchorages/splices developed in the tension zone of flexural r.c. members the deteriorated bond mechanism, as it is described by the local bond – slip law, results to a less stiff response and may lead to a premature and abrupt failure by affecting the deformation capacity of the member (Pantazopoulou et al. 2018). This is because corrosion and bond rely on the brittle concrete cover response in hoop tension, σ_h, as both mechanisms generate radial pressure to concrete, σ_r (Figure 1, i.e., the transverse equilibrium results to D_bσ_r = 2cσ_h where D_b = bar diameter and c = cover thickness). The negligible tensile deformation capacity of concrete upon attainment of its tensile cracking strength results in abrupt cover splitting associated with commensurate reduction of bond strength. The latter is further degraded by the unhindered ingress of corrosive agents resulting in ribs elimination. This dual effect is well described by the simple frictional analogue for estimating bond strength, f_b, as f_b = μ·σ_r = μ·2c/D_b·σ_h; the deterioration of both terms, the σ_r when σ_h exceeds the matrix tensile strength (σ_h = f_t) and the frictional coefficient μ that accounts on the impaired fib profile, results in poor bond resistance.

Figure 1:

Tensile stresses in concrete generated by bond and flexure.

By replacing conventional concrete with a ductile media, where strength is sustained in a wide range of deformation, bond mechanism becomes resilient against corrosion attack because the media provides restrain against cracking initiation and evolution. Such media may be an engineered cementitious composite of high volume, randomly dispersed short plastic fibers (known as ECC – a term coined by Victor Li 1998) where the dense fibers network arrests the splitting crack opening (Figure 1).

The behavior of corrosion process of steel bar reinforced ECC (R/ECC) has been studied only recently. Chen and Yang (2019) reported damage due to steel bar corrosion in ECC of increase number of cracks but of much lesser opening in comparison with the fewer and wider cracks in reference (brittle) mortar. They claimed that the ECC delays crack initiation whereas upon cracking, corrosion products accumulate inside the microcracks by forming a dense rust layer which reduces the ingress of aggressive ions to reach the surface of the uncorroded steel bar and by forming less expandable types of rust. Chen et al. (2018) reported that corrosion rate of steel bar in ECC is much reduced in comparison with the same mix without fibers due to high tensile ductility and damage tolerance of ECC material together with tight cracking within ECC cover. Qiu et al. (2020), by investigating the R/ECC corrosion process through X-ray, deduced that the matrix experiences longstanding growth of tiny and uniformly distributed cracks originated at pores in the vicinity of steel bar; on the contrary, cracking in plain concrete tends to occur locally at the ITZ between aggregate and cement paste. Kobayashi et al. (2018) concluded that the low water cement ratio in ECCs results in better corrosion protection because chloride entered from cracks slowly penetrates by convection or dispersion in the matrix and thus affects smaller region of the bar/matrix interface. Kobayashi et al. (2021) noticed that for the same amount of penetrated chloride the corrosion area and corrosion loss of rebar in ECC are both significantly lower than in concrete.

The flexural response of a reinforced concrete element has been analytically solved recently by Megalooikonomou et al. (2018), where the influence of bond mechanism is proved to dominate the flexural cracking evolution and deformation through bar slippage. In the case of corroded R/ECC elements as compared with counterparts made of ordinary concrete, Maalej et al. (2003) and Nguyen et al. (2018) noticed that the enhanced bond mechanism in the former case is responsible for smaller loss of strength and deformation capacity and for marginal affection of the flexural stiffness of R/ECC specimens. Aiming to explain the structural behavior of corroded R/ECC elements the necessary tool is the impaired local bond law. While there is abundance of data for local bond of corroded bars in concrete, the relevant studies where the media is an ECC are still limited. Hou et al. (2017, 2018 conducted standard pullout tests of pre-corroded steel bars embedded in intact ECC media (in this case the effect of bar corrosion evolution on the surrounding media was eliminated); with variables being the embedded length, bar diameter, and corrosion level, they concluded that the bond–slip law exhibits in comparison to standard concrete, a wider plateau at strength, higher residual resistance and increased bond toughness (the area enclosed by the bond curve and the axis of slip), all as a result of the confinement provided by the fibers. However, by adopting the standard pullout test, both the bond strength and the shape of the bond-slip curve are spuriously enhanced by the compressive stresses introduced by the support close to the bar loaded end. Such favorable conditions are not compatible with the stress state that occurs in anchorages developed in the shear span and in the tension zone of flexural members (Tastani and Pantazopoulou 2002). In Hou et al. (2019), ECC beams of varied anchorage length were corroded before testing; the results show that bond strength is unaltered up to corrosion level of 12% whereas the bond toughness in some cases is even increased.

The present study aims to enhance the available, but limited, experimental data on local bond of corroded steel bars developed in ECC media under adverse tensile stress conditions. To this end a total of eighteen short beams were built and tested. All were cast from the same cementitious mix: eight of them without fibers and the remaining ten with 2% per volume polypropylene fibers. Most specimens were subjected to artificial corrosion until cover cracking; the rest were left uncorroded for comparison. Finally, two different anchorage lengths were considered (extending into one shear span of the beams) in each of the aforementioned specimen groups. The experiment was meant to assess the restraining effect of the dense fiber network on the build-up of expansive rust, by measuring the degree of rebar corrosion that led to cracking of the concrete cover, and the consequent improvements in residual anchorage capacity, by measuring the mechanical response to four-point loading.

2 Experimental program

2.1 Materials and mechanical properties

The strain resilient cementitious composite (SRCC) used in the present study is a class of ECC, that follows the rule of combining fine-grained mixing materials with a high dosage of plastic fibers. The per weight composition was: one part Cement (CEM I 52.5N), three parts Greek calcareous fly ash (around 44% of SiO₂, Al₂O₃ and Fe₂O₃ combined, 40% of CaO, and 14% of free lime), 1.1 parts dried silica sand (100% passing through the 0.4 mm sieve) and 1.47 parts water [the water to binder ratio being 1.47/(1 + 3) = 0.37], with the addition of superplasticizer (SIKA, Viscocrete Ultra 200) at 7% of binder weight, and of polypropylene fibers at 2% per volume (fibers with diameter d_f = 25 μm, length ℓ_f = 12 mm, fracture stress and strain, respectively, f_fu = 400 MPa – ε_fu = 0.25, and density ρ_f = 0.91 gr/cm³, type ΤΜΙΧ-12 by ThracePlastics).

Previous studies of SRCC material in bending (Tastani et al. 2017) showed that, after first cracking, the deformation is concentrated into a limited number of cracks due to the fibers low stiffness (modulus of elasticity E_f ≈ 5 GPa) as compared with that of the cementitious matrix. Macroscopically, its flexural response follows a wide parabolic shape, sustaining adequate tensile strength up to high levels of deformation. The mechanical properties of this matrix (denoted by letter F, as fiber reinforced) as well as of the accompanied matrix without fibers (denoted by P, as plain matrix) in compression and in flexural tension were measured at the same age of the beam testing and the results are shown in Figure 2a and b. The data from flexural testing (Figure 2b) were back-analyzed by following the proposed procedure of Georgiou and Pantazopoulou (2016) aiming at defining the direct tension stress – strain law of the material (Figure 2c).

Figure 2:

SRCC mechanical properties: (a) compression (cubes of 100 mm), (b) flexural tension and (c) axial tension law of the SRCC material from inverse analysis of data from (b).

2.2 Beam specimens design

The experimental program comprised eighteen beams [width (b) × height (h) × clear span, 100 × 80 × 450, in mm] in four-point bending (shear span L_s = 180 mm), each reinforced with two steel bars of diameter D_b = 10 mm and nominal yielding stress f_y = 500 MPa, with bottom and side clear cover of c = 20 mm. The cross section effective depth is d = 55 mm (=h-c-0.5 D_b) and the reinforcement ratio is defined as ρ_l = A_s/(bd) = 2.85%. The shear span to effective depth ratio is L_s/d ≈ 3.3 (and L_s/h ≈ 2.3). The short anchorage length of either L_b = 5D_b or 10D_b was developed in the end region of the right shear span and in the tension zone of the cross section (Figure 3a and b). In the remaining segment of that shear span, L_s-L_b, the constant moment region and the ends, the reinforcing bars were covered by a sleeve aiming at bond elimination (Figure 3a). With this configuration the non-corroded bar tensile force F_s at the position x = L_s (where moment is maximum, M = QL_s/2, Q is the applied load and x is measured from the right support) is fully transmitted to the starting point of the corroded anchorage (x = L_b), as shown in Figure 3c. In the opposite shear span, the bars were fully bonded. Ten specimens were cast with SRCC and the remaining eight with the plain mix without fibers (to be used for reference). All specimens were cured in a laboratory chamber (95%RH, 20 °C) for 28 days and then left in room conditions for 4 months.

Figure 3:

Beam test: (a) detailing before casting, (b) geometry and setup and (c) regions of the reinforcing bar where bond action f_b is active.

2.3 Accelerated corrosion and results

In specimens intended for artificial corrosion, the bottom and sides were covered by impermeable resin (SIKADUR 330) with only the study region of L_b left uncovered, to be exposed to corrosion agents (Figure 4a); the end region of 15 mm was also uncovered, however the sleeve protects the bar against corrosion. Ten specimens were placed in a tank with aqueous solution of NaCl (3% per weight) and attached to an electrochemical corrosion cell, connected in parallel to a common electrical potential of 5 V, as shown in Figure 4b.

Figure 4:

Accelerated electrochemical corrosion: (a) specimen preparation and (b) detailing of the electrochemical cell.

The corrosion process lasted three months (a four-day cycle protocol was followed where one day of bottom surface wetting was followed by three days of drying). Measurements of electric current were taken every 24 h. The rate of steel mass loss Δm/Δt due to oxidation was calculated by following Faraday’s Law [i.e. Δm/Δt = A_mI/(zF) in gr/sec, where A_m = 55.87 gr is the atomic mass of iron, I is the current density in Amp, z = 2 is the valence of the reaction and F = 96486.7 Coulombs/mole is the Faraday constant]. The total mass loss ΔM_s (in grs) was estimated by integration of the current I passed through both bars of each specimen, where the time interval is 24 h. Assuming that mass loss occurred uniformly along the exposed length ℓ_exp of both bars in each specimen [ℓ_exp = 2·(a + L_b), see Figure 3] and using the definition of iron density (ρ_s = ΔM_s/ΔV_s = 7.87 gr/cm³) it follows that the mass loss, ΔM_s/M_s, is equal to the cross section area loss, ΔA_s/A_s, and approximately equal to twice the bar diameter loss, ΔD_b/D_b. The results of the diameter loss, ΔD_b/D_b and the naming conventions for the beam specimens are given in the first two columns of Table 1; for example in cP10D-2, c denotes corroded specimen (absence of letter c denotes non-corroded specimen), P is for plain matrix (or F for SRCC matrix), 10D refers to the bonded length (or 5D), -2 the coupon number of the subgroup.

Table 1:

Experimental results (corrosion measurement and mechanical testing) and bond estimation in the study region L_b.

Specimen ID^a	ΔD_b/D_b (%)	Q _max (kN)	Δ_max (mm)	Δ_90%Qmax (mm)	μ _Δ	Failure mode	Bond estimation f_b at Q_max (MPa)
P5D-3	–	16.3	2.1	2.2	1	Bond	10.5 (f_b,max)
P5D-4	–	14.1	2.3	2.4	1	Bond	9.1 (f_b,max)
cP5D-1	8.7	18.6	2.5	2.8	1.1	Bond	12.0 (f_b,max)
cP5D-2	10.5	19.4	2.1	2.2	1	Bond	12.5 (f_b,max)
F5D-1	–	19.2	3.8	6.2	1.6	Bond	10.3 (f_b,max)
F5D-5	–	17.2	3.3	4.9	1.5	Bond	8.9 (f_b,max)
cF5D-2	12.5	24.9	5.1	5.6	1.1	Bond	14.2 (f_b,max)
cF5D-3	17.9	30	3.4	3.7	1.1	Bond	17.8 (f_b^max)
cF5D-4	12.5	31.3	4.8	6.8	1.4	Shear	18.7 (f_b)
P10D-3	–	17.1	1.5	1.5	1	Shear	5.5 (f_b)
P10D-4	–	14.9	1.5	1.5	1	Shear	4.8 (f_b)
cP10D-1	7.4	16	1.7	1.7	1	Bond	5.2 (f_b,max)
cP10D-2	7.7	13	1.5	1.6	1	Bond	4.2 (f_b,max)
F10D-2	–	30.5	4.2	6.7	1.6	Shear	9.1 (f_b)
F10D-5	–	33.7	4.6	7.6	1.7	Shear	10.2 (f_b)
cF10D-1	5.3	26.5	4	9.3	2.3	Shear	7.7 (f_b)
cF10D-3	4.9	24.4	3.6	8.4	2.3	Shear	6.9 (f_b)
cF10D-4	6.6	31.2	4.4	9.9	2.3	Shear	9.3 (f_b)

^ac for corroded specimens; P for plain and F for SRCC matrix; iD refers to the bond length, -j the coupon of the subgroup, μ_Δ = Δ_90%Qmax/Δ_max.

Figure 5 shows a comparative diagram of the term ΔD_b/D_b for the ten corroded specimens along with the cracking state of the exposed region at the end of the corrosion process due to rust volume expansion of the rebar: all specimens developed splitting cracks (side and/or bottom) along the exposed length. The bar diameter loss, ΔD_b/D_b, in short anchorage specimens (L_b = 5D_b) was 9.6% for plain matrix (cP5D) and 14.3% for SRCC matrix (cF5D); in long anchorage specimens (L_b = 10D_b), it was 7.5% for plain matrix (cP10D) and 5.5% for SRCC matrix (cF10D), respectively. For the same duration of corrosion, the long anchorage specimens apparently showed lower corrosion levels than the short anchorage ones. The favorable effect of the strain resilient matrix in restraining cracking and thus delaying the corrosion process (hence lowering ΔD_b/D_b as compared to the plain matrix) was only observed in long anchorage specimens.

Figure 5:

Cracking along the anchorage length of reinforcing bars due to accelerated corrosion process.

3 Experimental behaviour of beams

Mechanical testing of the beam specimens was conducted after the end of the corrosion process by following a displacement-control protocol of 0.5 mm/min. The maximum applied load and the associated mid-span deflection (Q_max, Δ_max), the deflection at 10% loss of strength, Δ_90%Qmax, and the failure mode of the specimens are also summarized in Table 1. Failure denoted as “bond” manifests anchorage failure in the study region whereas “shear” denotes shear failure in the opposite shear span, where full bond conditions were ensured for the reinforcing bars (see f_b^l in Figure 3b). Figure 6 presents snapshots at the end of each experiment, where the mode of failure is also denoted.

Figure 6:
Cracking failure patterns of all non-corroded and corroded beam specimens.

Figure 6:

Cracking failure patterns of all non-corroded and corroded beam specimens.

More specifically, by “bond” failed the following specimens: in the case of the non-corroded specimens P5D failure occurred due to the development of splitting cracking along the L_b (Figure 6a.i-ii), for the cP5D new splitting or opening of the existing cracks occurred (Figure 6b.i-ii), for the cP10D opening of the existing cracks occurred (Figure 6f.i-ii), for the non-corroded F5D bar slippage occurred with the development of a transverse crack at the initiation point of the anchorage without any sight of splitting along the L_b (Figure 6c.i-ii) and for the corroded cF5D opening of the existing cracks occurred (Figure 6d.i-iii).

By “shear” failed one specimen cF5D (Figure 6d.iii), all of P10D (Figure 6e.i-ii), F10D (Figure 6g.i-ii) and cF10D (Figure 6h.i-iii). All F10D and cF10D specimens developed similar behavior: at least three transverse cracks occurred into the left shear span of similar spacing, which at advanced state of deformation were connected with inclined cracking; the most remote crack was located at about 80 mm from the left free side. On the other hand, the P10D specimens developed two transverse cracks, the most remote crack being located at about 60 mm from the left side; at peak load abrupt shear failure occurred as it was highlighted by an inclined crack extended from the support to the loading point.

Figure 7 demonstrates experimental curves of applied load, Q, versus mid-span deflection, Δ. The green and black curves correspond to the non-corroded and the corroded specimens, respectively; in the latter case, the curve thickness signifies the corrosion level. By comparing the non-corroded F- and P- short anchorage specimens that exhibited bond failure (green curves of F5D, P5D in Figure 7), it becomes apparent that the addition of fibers resulted in an impressive increase of displacement capacity and a marginal increase of the load-currying capacity: on average, F-specimens reached the peak point at deflection 3.5 mm and load 18 kN, that was followed by a mild descending response, whereas P-specimens failed abruptly at deflection 2 mm and load 15 kN. These results show an increase in deflection and strength by 75% and 20%, respectively, when the bending response of the two matrices (Figure 2b) differ in deflection at peak by around 400% but are of similar strength of around 1.5 kN.

Figure 7:

Load–mid-span deflection responses of control (green) and corroded (black) specimens: for comparison the scale of the relevant axes is the same.

The corroded short anchorage specimens cF5D and cP5D that failed by bond developed higher load currying capacity as compared to their controls (increase of around 50% and 25%, respectively). Considering that their bar diameter loss (on average) is 14% and 9.6% respectively, the cF-specimens of higher degree of bar loss developed higher bond strength than the cP-specimens. This may be attributed to the ongoing maturing of the cementitious matrix as a result of the presence of moisture during the simulated corrosion that favors the bond conditions. By comparing the F5D and cF5D specimens, it seems that corrosion induced a steeper post-strength response; the accommodated rust around the bar, even if it is confined, after a certain magnitude of bar slip deteriorate the frictional coefficient resulting in steeper drop of load.

The long anchorage specimens F10D and cF10D failed by shear in the left span. Corrosion degree (in average 5.5%) slightly affected the response in terms of initial stiffness (reduction), strength (reduction) and deformation capacity (increase). On the other hand by comparing the P10D and cP10D specimens, corrosion (on average 7.5%) altered the mode of failure from shear in the left span to bond in the right span and induced slight decrease of stiffness and strength.

4 Data analysis and discussion

4.1 Flexure

Aiming to explain the ductile response of the F5D specimens (region extending from the peak point (Coronelli et al. 2019; MacGregor and Wight 2005) to that of marginal strength loss of 10%, the deformation was increased from 3.5 mm to 5.6 mm resulting in displacement ductility of μ_Δ = Δ_90%Qmax/Δ_max = 1.6), a cross section flexural analysis is performed at bar yielding strain (ε_sy = 0.0025) for the mechanical properties of matrices F and P as shown in Figure 2a and c. More specifically, the compressive stress – strain response up to the peak point is approximated by Hognestad’s parabola (MacGregor and Wight 2005) as σ = 0.85f_co[2ε/ε_co − (ε/ε_co)²], with ε_co|F = 0.008 − f_co|F = 45 MPa, ε_co|P = 0.008 − f_co|P = 80 MPa; in tension for the matrix F, a bilinear response is adopted with milestone values at cracking and at ultimate the average among the inverse analysis data of the two types of prisms (Figure 2c), with ε_t,cr|F = 0.00015 − f_t,cr|F = 2 MPa and ε_t,u|F = 0.02 − f_t,u|F = 2.5 MPa, and for the matrix P, a linear of ε_t,cr|P = 0.00015 − f_t,cr|P = 4.5 MPa. The theoretical yielding loads are calculated as Q_y|F ≈ 40 kN and Q_y|P ≈ 38 kN (at compressive strain of the extreme cross section fiber of ε_c|F = 0.0063 and ε_c|P = 0.0033, both lower than ε_co). By comparing Q_y values with the peak loads of Figure 7, none of the uncorroded specimens reached the yielding load. So, the ductile response of F-specimens is not the result of the bar yielding, but rather is attributed to the controlled sliding of the bar with respect to the SRCC matrix, due to the confinement provided by the PP fibers.

4.2 Shear

For the non-corroded P10D and F10D specimens, while the initially expected mode of failure by bond did not occur, their comparison highlights the contribution of fibers in shear strength, V_shear, as V_shear = V_c + V_f where V_c is the contribution of the matric (plain or fiber reinforced) and longitudinal reinforcement and V_f is the contribution of the fibers. From Table 1, by considering the average value among the same subgroup, it is V_shear^P = V_c^P = Q_max^P10D/2 = 16/2 = 8 kN [which is lower than the estimated by EN1992-1-1 (2004) as V_c = C_Rd,c·k· (100·ρ_l·f_c)^1/3·b·d = 0.18·2·(100·0.02·80)^1/3·100·55 = 10.8 kN, on which the design of the plain matrix beams was based aiming at precluding shear failure] and V_shear^F = V_c^F + V_f = Q_max^F10D/2 = 32.1/2 ≈ 16 kN. Hence, the fibers increase the shear strength by 100%. Georgiou and Pantazopoulou (2017) have proposed Equation (1) for the prediction of the shear capacity of reinforced ECC beams without stirrups:

(1)VshearF=π·Db·nb·fb,max·(1/3y−c+2/3h)+b·ft·y·(4h−y)/(6Ls)

where n_b is the number of tensile bars, the bond strength is f_b,max =(2μ/π)·(c/D_b)·f_t where the matrix tensile strength is taken as f_t = 2.5 MPa (see Section 4.1) and the frictional coefficient is μ = 0.9 hence f_b,max=(2·0.9/π)·(20/10)·2.5 ≈ 2.9 MPa, y is the height of the tensile depth of the cross section given as y = 0.6h = 0.6·80 = 48 mm, h and b are the section height and width, and c is the clear cover. The implementation of Equation (1) by using data of the present study results to V_shear^F = π·10·2·2.9(1/3·48–20 + 2/3·80) + 100·2.5·48(4·80–48)/(6·180)≈12 kN. For all F-specimens that failed by shear in the left shear span (non-corroded region), this estimation, when it is compared with Q_max/2 (Table 1), approximates only specimens cF10D-1 and cF10D-3; for the rest of them (i.e. F10D-2, F10D-5, cF5D-4 and cF10D-4) V_shear^F is lower by around 3 kN. An improvement of the frictional coefficient μ from 0.9 to 1.2 as a consequence of the higher fracture toughness of the ECC as compared with the plain matrix results in V_shear ≈15 kN, which is very close to the experimental values.

In general, the shear failure precedes anchorage failure when bond strength both in L_b and in the left span is stronger than the shear strength. Failure in the left shear span implies that the shear strength V_shear is lower than the transverse force V_anch associated with the anchorage capacity in the study region L_b where the bond stress f_b^r attains the strength f_b,max. The analysis of the data for the definition of the bond strength is presented in Section 4.3.

4.3 Bond

For the estimation of the bond stress f_b at peak load Q_max, that is equal to the bond strength f_b,max only for those specimens failed in the right shear span where L_b is developed, the implementation of inverse cross section analysis at peak load Q_max results to the associated bar strain ε_s (lower than the yielding limit, ε_sy = 0.0025) at the critical cross section (region of constant moment where the reinforcement cross section is intact due to sleeve). Given that the bar tensile force F_s is fully transmitted to the initiation point of anchorage due to the presence of the sleeve (Figure 3a and b), the attained bond stress value may be approximated by assuming uniform distribution along L_b through equilibrium as F_s = πD_b²/4·E_sε_s = πD_bL_bf_b, where E_s = 200 GPa. For specimens that failed by bond in the study region then f_b =f_b,max. Values of bond are shown in Table 1.

Even though the values of f_b,max for all control specimens with L_b = 5D_b (5D) are of similar magnitude (∼9.5 MPa on average), the bond toughness of the F-specimens is higher than of the P-type as is evident by comparing magnitudes of deflection ductility μ_Δ (1.5 and 1, respectively). The associated corroded specimens of both matrices developed higher bond strength than their controls; interesting is that the F-specimens that suffered higher degree of bar loss developed higher bond strength than the P-specimens. By comparing the F5D and cF5D specimens, it seems that corrosion induced a steeper post-strength response (reduction of µ_Δ form 1.5 to 1.1); the accommodated rust around the bar, even if it is confined, after a certain magnitude of bar slip deteriorates the frictional coefficient resulting in almost instant drop of load.

For the long non-corroded anchorages P10D and F10D, even if these beams failed by shear in the opposite shear span, the calculated bond stress values of F-specimens are double of P- specimens (9.5 MPa and 5 MPa, respectively) highlighting the beneficial effect the fibers have on bond mechanism. Also, the fact that the F10D bond stress is identical to f_b,max attained by the F5D specimens (9.5 MPa on average, also identical values for μ_Δ) denotes that the interface attained its capacity but its resilience prevailed against the shear mechanism. The shear-failed cF10D developed an average bond stress of 8 MPa, higher than the bond-failed cP10D (4.5 MPa, in average) and slightly lower than the F10D but of higher ductility.

The preceding discussion of the structural response of non-corroded and corroded steel anchorage ECC beams sums up as per the resilience of the composite matrix and its associated beneficial effect on bond mechanism even in case of reinforcement corrosion. The experimental results may be useful to practitioners in case of patch repair of inadequate anchorages/splices that suffer bar diameter loss due to corrosion of less than 10%.

5 Conclusions

The present experimental program aims at investigating the bond mechanism of corroded steel anchorages (short and long) embedded in a strain resilient cementitious matrix with 2% per volume PP fibers; for comparison associated specimens were cast with plain matrix. For common period of exposure to electrochemical corrosion, a strain resilient matrix as compared with a plain matrix can delay metal consumption because the fibers provide restrain against crack initiation and widening; this is evidenced by the corrosion results of the longer anchorages. For both matrices, the presence of moisture during the corrosion process resulted in increased local bond strength of the corroded specimens as compared to their controls (for L_b = 5D_b). For the short anchorages, while their control bond strength magnitudes are similar (∼10 MPa), the most corroded F-type specimens developed higher bond strength than the P-type. The post-strength response of cF5D tests failed by bond was affected by the corrosion resulting in a steeper drop of the load-carrying capacity. For the long anchorage specimens that seem to prematurely fail by shear, the data analysis highlighted that the bar-to-ECC matrix interface resilience prevailed against the more brittle shear mechanism. These experimental results may be useful to practitioners in case of patch repair of inadequate anchorages/splices that suffer bar diameter loss due to corrosion of less than 10%.

Corresponding author: Sousana Tastani, Department of Civil Engineering, Democritus University of Thrace, Xanthi67100, Greece, E-mail: stastani@civil.duth.gr

Acknowledgments

The experimental program was done without funding in the Lab. of Building Materials of Democritus University of Thrace, Greece, in the context of the diploma thesis of undergraduate student Dakidis Ch., whose contribution is greatly appreciated. Materials grant: superplasticizer from SIKA HELLAS, fibers from Thrace Plastics Co S.A., fly ash from DEI, Ptolemais, Greece.

Author contributions: The author has accepted responsibility for the entire content of this submitted manuscript and approved submission.
Research funding: None declared.
Conflicts of interest: The author declares no conflicts of interest regarding this article.

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Received: 2022-07-20

Accepted: 2022-12-03

Published Online: 2023-02-09

Published in Print: 2023-04-25

This work is licensed under the Creative Commons Attribution 4.0 International License.

Articles in the same Issue

https://doi.org/10.1515/corrrev-2022-0068

Keywords for this article

bond; cementitious composites; steel reinforcement corrosion

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