Engineering Structures 32 (2010) 2387–2393
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Accuracy of design code expressions for estimating longitudinal shear strength of strengthening concrete overlays E.N.B.S. Júlio a,∗ , D. Dias-da-Costa a,1 , F.A.B. Branco b , J.M.V. Alfaiate b a
ISISE, Department of Civil Engineering, Faculty of Sciences and Technology, University of Coimbra, Portugal
b
ICIST, Department of Civil Engineering, Instituto Superior Técnico, Technical University of Lisbon, Portugal
article
info
Article history: Received 27 November 2009 Received in revised form 9 February 2010 Accepted 1 April 2010 Available online 7 May 2010 Keywords: Interface Shear strength Design codes Concrete overlay Strengthening
abstract Strengthening operations of RC structures often imply the enlargement of the original cross-sections. In these cases, the shear strength of the resultant concrete-to-concrete interfaces is crucial in assuring the monolithic behavior of the strengthened members. Most concrete codes present design expressions for estimating the shear strength between concrete layers based on the shear–friction theory. However this was formulated for precast members with cast-in-place parts. Aiming to analyze the accuracy of code expressions in the strengthening situation, a research project was conducted. First, an experimental study was performed to evaluate the shear strength between a sandblasted concrete substrate and a concrete overlay, for different amounts of transverse reinforcement at the interface. Then, a numerical study was undertaken to enlarge the results’ range. Finally, these results were compared with those given by design codes and conclusions were drawn. It should be highlighted that: (a) the values of shear strength given by design code expressions are significantly different from each other; (b) in two cases, for low reinforcement ratios, these are not safe; and (c) in most cases, for high reinforcement ratios, these are too conservative. © 2010 Elsevier Ltd. All rights reserved.
1. Introduction
2. Research significance
Reinforced concrete (RC) bridge decks and building slabs are often strengthened by adding a concrete overlay. RC jacketing of columns and beams is also commonly used in strengthening operations. In these situations, concrete-to-concrete interfaces play an important role in assuming the monolithic behavior of the resulting composite RC members. Previous studies by the authors were focused on the influence of different parameters on the bond strength of old-to-new concrete interfaces: roughness of the substrate surface [1–3], use of epoxy-based bonding agents [4], and compressive strength of added concrete [5]. The present paper describes a subsequent study performed to analyze the influence of added reinforcement crossing the interface, based on and complementing the preceding work, but also aiming to verify if design codes are applicable in estimating the shear strength at the interface between the existing concrete member and the concrete overlay.
Design code expressions are based on the shear–friction theory [6–9], formulated for precast members with cast-in-place parts. Nevertheless, there are differences between precast/in situ composite structures and old structures being strengthened by overlays. In the first case, the reinforcement crossing the interface is placed before casting the original member and the substrate surface is left naturally rough or is roughened while fresh. In the second case, the reinforcement is usually epoxy-bonded to the hardened concrete member and the substrate surface is roughened with specific techniques for hardened concrete. Therefore, the authors considered it important to verify if design code expressions are also applicable in strengthening situations.
∗
Corresponding author. Tel.: +351 239 797 258; fax: +351 239 797 259. E-mail addresses:
[email protected] (E.N.B.S. Júlio),
[email protected] (D. Dias-da-Costa). 1 Tel.: +351 239 797 256. 0141-0296/$ – see front matter © 2010 Elsevier Ltd. All rights reserved. doi:10.1016/j.engstruct.2010.04.013
3. Reinforcement crossing the interface The design method to assess the shear strength of concrete-toconcrete interfaces has changed throughout the years. Nowadays the majority of design codes have adopted expressions based on the shear–friction theory [6]. According to this, the shear strength develops by friction between both concrete layers, as shown in Eq. (1):
νn = ρ · fy · µ
(1)
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Symbols
α αda µ νn ρ σn τcoh φ ψ c0 C1 , C2 fc fy GIIF k ks
bond-slip shape parameter in MC 90 dowel action coefficient friction coefficient shear strength by unit of area ratio between the area of the reinforcement crossing the interface and the area of the interface external normal force across the interface cohesion due to aggregate interlock internal friction angle dilatancy angle cohesion parameters in Walraven’s expression compressive strength reinforcement yielding stress mode-II fracture energy parameter in Loov’s expression elastic stiffness of the joint
where νn is the shear strength by unit of area, ρ is the ratio between the area of the reinforcement crossing the interface and the area of the interface surface; fy is the reinforcement yielding stress; and µ is the friction coefficient which assumes the following values: 1.7 for monolithic concrete, 1.4 for joints with the surface roughened artificially and 0.8 to 1.0 for current construction joints and steel–concrete interfaces. An important improvement was published by Mattock and Hawkins [7], today known as the modified shear–friction theory, given by Eq. (2):
νn = 1.38 + 0.8 · ρ · fy + σn
(2)
where σn is the effect of the external normal force across the interface. The following differences have to be highlighted in relation to the Birkeland and Birkeland expression: the friction coefficient assumes the value 0.8; normal stresses at the interface are not only due to tension of the reinforcement crossing the interface; and, for the first time, cohesion of the interface (understood as aggregate interlock) is also considered, assuming the value 1.38 MPa. Another important contribution was given by Loov [8], expressed as Eq. (3)
νn fc
s =k
ρ · fy + σn
(3)
fc
where, for the first time, the concrete compressive strength, fc , is taken into account, being k a constant equal to 0.5, for initially uncracked interfaces. Many other contributions followed, e.g. by Walraven et al. [9], who performed a statistical analysis of push-off test results and suggested an expression assuming perfectly spherical aggregates, given by Eq. (4):
νn = C1 ρ · fy
C2
(4)
where C1 = 0.822fc (MPa)0.406 and C2 = 0.159fc (MPa)0.303 , having the remaining symbols the same meaning as before. Finally, Randl [10] proposed a design expression that separates and explicitly incorporates the three influencing parameters (cohesion, friction and dowel action), as shown in Eq. (5):
νn = τcoh + µ · σn + αda · ρ ·
p
fc · fy
(5)
where τcoh is the cohesion due to aggregate interlock and αda is a coefficient to take into account the dowel action.
From the examples given before it can be understood that, although based on the shear–friction theory, quite different design expressions have been proposed to assess the shear strength at concrete-to-concrete interfaces. For this reason, it is not surprising that in different design codes, from different countries and edited in different years, also quite different design expressions are presented. In the present study, the following design codes were considered: Portuguese Code (REBAP) [11], Euro Code 2 (EC 2) [12], Model Code 90 (MC 90) [13], Canadian Code (CSA) [14], USA Code (ACI 318) [15], and British Code (BS 8110) [16]. Table 2 presents the expressions to calculate the design value of the shear strength of the interface, according to each of the selected codes, for the conditions considered in the study herein described. The maximum values assumed by these codes are also given. It must be clarified that: (a) safety coefficients were not applied; (b) transverse reinforcement was assumed to be applied at a 90° angle to the interface; (c) the interface was assumed to be clean and rough; (d) a characteristic compressive strength of 35 MPa was considered; and (e) no external normal stress was present. 4. Experimental research The experimental work described in this paper was defined to: (a) determine the concrete-to-concrete interface debonding stress (at the instant adhesion is lost), for different percentages of reinforcement crossing the interface; (b) analyze the corresponding behavior, after debonding of the interface, and determine the interface shear strength; (c) verify the difference between having the reinforcement placed before casting the substrate concrete and having it inserted into the hardened concrete substrate; (d) for this second situation, analyze the efficiency of two commercial products used to anchor the steel connectors; and (e) compare test results with values determined according to design codes. The experimental study was planned based on conclusions drawn from previous studies conducted by the authors and previously mentioned. It is well known that the roughness of the substrate highly influences the strength of the interface. Therefore, the choice of an adequate method to prepare the substrate surface is important. Although hydrodemolition is considered to be the best surface treatment [17,18], since sandblasting is also one of the most efficient techniques and the best from those previously adopted by the authors [1], this was used in all situations, excepting one (left without treatment), considered as reference. In some strengthening situations, a bonding agent is also used to increase the bond strength of old-to-new concrete interfaces. However, this was not adopted in the present study because in a previous study [4] the authors proved that the application of an epoxy-based bonding agent on the substrate surface does not improve the bond strength of the interface if sandblasting, or another method that adequately increases the roughness of the substrate surface, is chosen. Generally, the mechanical properties of the added concrete layer and of the existing concrete substrate are considerably different. Previously [5] the authors have concluded that the compressive strength of the added concrete has a significant influence on concrete-to-concrete bond strength. Furthermore, it was demonstrated that high strength concrete shows advantages for carrying out repairing and strengthening techniques involving adding new concrete to an existing concrete substrate. However, the influence of this parameter is not taken into account in any design expressions since only the weakest concrete is considered. In the present study, it was decided to adopt similar mechanical properties for both the substrate and added layer. The push-off test was selected to perform the study herein described (Fig. 1). The adopted push-off specimens are antisymmetrical, made of two identical ‘‘L’’ shape halves. The geometry
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127
127
127
19
19 58
19 58
127
2389
69
138
69
127 19
58 19
58 19
127
127
127
37
180
37
37
180
37
37
180
37
Fig. 1. The adopted push-off specimen (dimensions in mm) with: (a) two; (b) four; and (c) six steel connectors. Table 1 Experimental program: considered situations and characteristics of the tested specimens. Situation
Surface treatment
Anchoring system
Number of steel connectors
(1) (2) (3) (4) (5) (6) (7)
No treatment
– – Manufacturer 1 Manufacturer 1 Manufacturer 1 Manufacturer 2 Previously embedded
0 0 2 4 6 6 6
Sandblasted surface
Table 2 Shear strength of the interface according to different codes for the conditions considered in the experimental study. Codes
REBAP EC 2 MC 90 CSA ACI 318 BS 8110
Shear strength of the interface Design expression (MPa)
Maximum value (MPa)
νn νn νn νn νn νn
– 11.1 10.8 10.8 5.5 –
= ρ fy = 1.28 + 0.7ρ fy = 1.49 + 0.9ρ fy = 0.25 + 0.6ρ fy = 0.6ρ fy = 0.8ρ fy
of the envelope is a 254 × 546 × 127 mm3 prism, adopted from Hofbeck et al. [19]. Each half is reinforced with nine S400 steel bars with 10 mm diameter and eight S400 steel stirrups with 6 mm diameter. For the reinforcement crossing the interface, it was also adopted S400 steel bars with 6 mm diameter and an average yielding strength of 443 MPa and a tensile strength of 553 MPa. Seven situations were considered and are summarized in Table 1. In situation (1), the substrate surface was cast against steel formwork, without roughness treatment and without steel connectors, considered as reference. For the remaining situations, based on previous studies, the interface surface was prepared by sandblasting. The number of S400 steel connectors with 6 mm diameter and the commercial product used to bond these to the concrete substrate varied. The following situations were considered: (2) 0 steel connectors; (3) 2 steel connectors, anchored with an epoxy resin by manufacturer 1 (Fig. 1(a)); (4) 4 steel connectors, anchored with an epoxy resin by manufacturer 1 (Fig. 1(b)); (5) 6 steel connectors, anchored with an epoxy resin by manufacturer 1 (Fig. 1(c)); (6) 6 steel connectors, anchored with an epoxy resin by manufacturer 2 (Fig. 1(c)); and (7) 6 steel connectors, previously embedded in the concrete substrate (Fig. 1(c)). All parameters that may have influence on results, apart from the reinforcement crossing the interface, were kept constant. The same concrete mix was adopted both for the substrate and for
Fig. 2. Different steps of production of push-off specimens: (1) assembling substrate rebars; (2) substrate casting; (3) application of connectors; (4) assembling added half rebars; (5) added half casting.
the added layer. An average compressive strength of 43 MPa was obtained at 28 days of age. The age of the substrate concrete and the age of the added concrete were settled as 16 and 4 weeks, respectively, at the time of test. The only exceptions were the relative humidity and temperature. However, these two parameters were exactly the same for all substrate halves and all added halves of the seven situations considered. Fig. 2 illustrates the relevant steps of specimens production. For each of the seven situations considered, five specimens were produced according to the following procedures: (1) assembling of the reinforcing bars and stirrups of the substrate half; (2) casting of the substrate half; (3) execution of holes, anchoring of steel connectors using an epoxy resin and roughness increase of substrate surface by sandblasting; (4) assembling of the reinforcing bars and stirrups of the added half and bonding of two polystyrene plates on the original half to produce a 19 mm gap; (5) casting of the added half. Push-off tests were performed on a 500 tf universal testing machine with displacement control (Fig. 3). A load cell, TML CLC100A, and two displacement transducers, TML CDP 25, connected to a data logger, TML TDS 602, were used in addition to the control system of the universal testing machine. This redundant data acquisition had the purpose of detecting possible errors and increasing results accuracy. 5. Discussion of experimental results Fig. 4 shows the load–deformation curves of one randomly chosen push-off specimen tested, for each of the five situations considered with reinforcement crossing the interface surface. Qualitatively, the 35 tested specimens presented the same behavior: for a constant increase of deformation with time, a rapid increase of load was observed, until debonding of the interface surface occurred. For this instant, a sudden loss of both strength and stiffness was observed. After debonding, for the same constant increase of deformation with time, a slow increase of load
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5.0
Fig. 3. (a) Push-off test; (b) instant of interface debonding; (c) end of test.
Interface debonding st ress [M Pa ]
4.5 4.0
3. 1
3. 7
3. 8
S3 S4 S5 Test ed Sit uations
S6
3. 2
3. 4
3. 9
3.5 3.0 2.5 2.0
1. 8
1.5 1.0 0.5 0.0 S1
S2
S7
Fig. 5. Average interface debonding stress and corresponding interval at 75% confidence (MPa).
Fig. 4. Load (kN) versus displacement (mm) curves for one specimen of situations (3)–(7).
was registered. Higher loads were reached with the increase of deformation, after debonding of the interface surface, for higher percentages of steel connectors applied perpendicular to the interface surface. In Figs. 5 and 6, for each of the situations considered, the average debonding stress of the interface and the corresponding average maximum shear stress (reached after debonding) are given. Results of push-off tests of specimens without steel connectors, situations (1) and (2), substantiate the conclusion, drawn in a previous study [1], that, by preparing the substrate surface with sandblasting, a higher average debonding stress of the interface is achieved. Nevertheless, the registered difference is lower than the corresponding value obtained with the slant shear tests performed before [1]. Comparing situations (3) to (5), with the substrate surface treated with sandblasting and with, respectively, 2, 4 and 6 steel connectors applied crossing the interface and anchored with an epoxy resin by manufacturer 1, it was noticed that: (a) the average debonding stress is approximately the same with 2, 4 or 6 steel connectors; and (b) the behavior after debonding of the interface varies significantly with the number of steel connectors crossing the interface surface. In what concerns the two commercial bonding agents used to anchor six steel connectors, situations (5) and (6), results are not significantly different. Furthermore, these results are also similar to those of the corresponding situation with six steel connectors previously embedded, situation (7), both in terms of debonding stress and of behavior after interface debonding. 6. Numerical study and comparison with experimental results and codes Aiming to better analyze the design code expressions, a numerical study was conducted to broaden results. A finite-
element package [20–22] developed for bidimensional problems addressing the simulation of structural interfaces and quasi-brittle fracture, by means of discrete strong discontinuities, was adopted. First, the numerical model was calibrated to reproduce the obtained experimental results. Then, with the calibrated model, situations with higher values of ρ were numerically simulated. In this study, concrete was assumed to have a linear elastic and perfectly plastic behavior under compression, limited by an average compressive strength of 43 MPa, experimentally assessed. Under tensile stresses, linear elastic behavior was also assumed. Since it was considered that the provided steel reinforcement absorbs these stresses, a limiting tensile strength was not defined, thus avoiding the possibility of concrete cracking. Concrete Young’s modulus and other material properties not experimentally obtained were computed from EC 2 [12]. The progressive tensioning of the steel connectors induces compressive stresses along the interface, leading to an increased shear strength capacity. Therefore, the rheological behavior of the interface was assumed to follow a plasticity model with a Mohr–Coulomb friction law yield surface, without tensile cap [23]. For the steel connectors, a multilinear constitutive law adjusted to experimental results was considered. Experimental assessment of the bond stress–slip relation between steel connectors and concrete was not performed. For that reason, MC 90 [13] provisions were adopted, assuming confined concrete and good bond conditions. In what concerns the boundary conditions, first the numerical model was assumed restrained at both ends to simulate the effect of the bearing plates of the testing machine. Then, unrestrained conditions were considered, as if roller bearings have been used in the experimental part. This revealed that, without roller bearings, the top and bottom ends of the specimen almost completely detach from the plates and that the horizontal restraining force is approximately 20% of the vertical load. For this reason, it was decided to simulate all situations considering roller bearings. In the following section, only these results are analyzed and compared with theoretical values given by the six selected design codes to draw conclusions. The numerical model was composed of plane stress bilinear finite elements for concrete. At the structural joint level, two-node zero-thickness finite elements were used to connect the bilinear elements from each side of the joint. Crossing the interface, linear truss elements were adopted to simulate the connectors. These were subsequently connected to the concrete bilinear elements using zero-thickness finite elements and assuming the mentioned
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Maximum shear stress after debonding [MPa]
5. 0 4.5 4.0
3. 5
3. 7
3. 8
3.5 3.0
2. 6
2.5 2.0 1.5
1. 1
1.0 0.5 0.0 S1
S2
S3
S4
S5
S6
S7
Tested Situations
Fig. 7. Load (kN) versus displacement (mm) curve superposed with the experimental envelope, for 2, 4 and 6 steel connectors anchored with an epoxy resin by Hilti.
Fig. 6. Average maximum shear stress at the interface, after debonding, and corresponding interval at 75% confidence (MPa).
steel–concrete bond law. All finite elements were numerically integrated with a 2 × 2 Gaussian scheme, whereas a 2-point Newton–Cotes rule was applied for the zero-thickness finite elements. The numerical solution was obtained by using the arclength method, constraining the evolution of the shear opening of the joint. Several material parameters have been identified as relevant in the described constitutive models: initial elastic stiffness of the joint, ks ; internal friction angle, φ ; dilatancy angle, ψ ; cohesion, c0 ; mode-II fracture energy, GIIF ; and bond-slip shape between connectors and concrete measured by α , according to MC 90 [13]. A preliminary study allowed accurately defining loading and boundary conditions. Moreover, the role of each parameter, at each stage of the load versus displacement, was evidenced. From this knowledge, a strategy was defined to evaluate each parameter: (1) ks was chosen to obtain the same initial stiffness as the plain push-off specimens; (2) φ was evaluated to ensure similar residual strength, followed by a sensitivity analysis in relation to ψ , evaluating the hardening rate; (3) peak load was analyzed by means of c0 and GIIF , because all remaining parameters were already known at this stage; and (4) the softening part of the diagram and minimum value after peak load were studied by varying GIIF and α . According to the defined procedure, the following values were adopted for steel connectors anchored with an epoxy resin by manufacturer 1: ks = 18 MPa/mm; tan φ = 0.95; tan ψ = 0.18; c0 = 3.8 MPa; GIIF = 3.0 N/mm; and α = 0.3. The corresponding load versus displacement curve is represented in Fig. 7, where good agreement can be observed for all test situations. However, in the region of the minimum load, a local plateau appears. This corresponds to the yielding of the steel connectors while cohesion c0 is still decreasing. Truss elements cannot simulate bending, which constitutes a limitation of the numerical model. Experimentally, failure occurs in a mixed mode, under traction and shear, in which traction seems to be the predominant mechanism which can be well simulated by truss elements. The adjusted numerical model was rerun for higher values of ρ , considering 2, 4 and 6 steel connectors with the following diameters: 8, 10 and 12 mm. Fig. 8 presents the corresponding load versus displacement curves. It is emphasized that the gap between the halves has been doubled to allow for the failure of the specimen. It can be concluded that the increase on the load capacity is proportional to the area of the steel connectors, except when failure occurs with the steel connector debonding. Moreover, with increasing steel percentage, the softening branch
Fig. 8. Load (kN) versus displacement (mm) curve for 2, 4 and 6 steel connectors with different diameters.
of the load versus displacement vanishes. For the range of tested diameters, it is observed tensile failure for 8 and 10 mm diameters; whereas for 12 mm, failure occurs with the connectors debonding. For higher strengths, crushing of the concrete surrounding the upper gap and, later, on surrounding the lower gap, is observed. Therefore, the increase on the strength of the specimen is limited by concrete compressive strength and bonding of steel connectors. Since this was not observed in the experimental tests conducted with 6 mm connectors, it is advisable to perform further tests with 12 mm connectors in order to validate the corresponding numerical results. Fig. 9 shows the relationship between the shear strength and the normal stress at the interface according to each one of the codes adopted, superimposed with results of the push-off tests and of the numerical study. It must be stated that inherent to the shear–friction theory is the existence of slip, therefore the shear strength of the interface is assumed as the average maximum shear stress at the interface after debonding. For low normal stresses, it can be observed that, apart from MC 90 and EC 2, all the selected codes represent a conservative solution. This can be explained since only these two codes and CSA consider cohesion, but in the latter the influence of this parameter is much lower. For high normal stresses, it can be observed that, apart from REBAP and MC 90, all the selected codes give too conservative values. In fact for the experimental/numerical results the shear strength is proportional to approximately 1.3 times the normal stress, whereas for the corresponding theoretical relationship the proportionality varies between 0.6 and 1.0.
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14 ex p. results (without roller bearings )
REBAP
num . results (without roller b earings)
12
num . results (wit h roller bearings )
Shear Strength [MPa]
10
BS 8110
8
MC 9 0 CS A
EC 2
6 ACI 318
4
2
0 0
2
4
6
8
10
12
14
Normal Stress (ρfy) [MPa] Fig. 9. Experimental/numerical versus analytical ratios ‘‘shear strength/normal stress’’, according to different codes.
7. Summary and conclusions Based on and complementing previous studies performed by the authors, the joint experimental/numerical research described in the present paper was conducted to assess the shear strength between an original concrete substrate and a new concrete layer with added reinforcement crossing the interface. Furthermore, the study described herein aimed to verify if design code expressions for shear at the interface between concretes cast at different times are applicable to the strengthening overlay situation as well as to the precast/in situ composite situation for which these were derived. Push-off tests were performed considering four different amounts of transverse reinforcement; two roughness levels of the substrate surface; and two commercial resins for anchoring the reinforcement. Numerical models were built to simulate percentages of reinforcement crossing the interface outside the range of those considered in tests. Results were compared with theoretical values given by six selected design codes. From an analysis of the experimental results it was possible to conclude that: (1) the reinforcement crossing the interface does not significantly increase the interface debonding stress; (2) the shear strength of the interface increases with the increase of reinforcement crossing the interface; (3) for low reinforcing ratios, the shear strength of the interface corresponds to the debonding stress; (4) for higher reinforcing ratios, the shear strength of the interface is not reached immediately on debonding but only after an important slip; (5) there is a difference of 6.6% to 8.3% between having the reinforcement placed before casting the substrate and having it inserted into hardened substrate; (6) results obtained with each of the two commercial epoxy resins used to anchor the steel connectors were only marginally different; and (7) higher shear strength of the interface is achieved with sandblasted surfaces than with surfaces cast against steel formwork, corroborating conclusions of previous studies by the authors.
In relation to design codes, it can be concluded that: (8) substantial differences are registered between the values given by each of the codes considered; (9) comparing experimental/numerical values with code expressions, it can be stated that the values given by EC 2 and MC 90 for low reinforcing ratios are not safe; (10) the ratio of the shear strength to normal stress (the friction coefficient) is approximately 1.3 for experimental results and numerical analysis data; significantly higher than the corresponding ratio for code expressions which vary between 0.6 and 1.0; and (11) due to this fact, for higher levels of normal stress the codes, with the exception of REBAP and MC 90, tend to predict the shear strength conservatively. Concerning the differences between the interface of precast/in situ composite elements (situation in which codes are focused) and the interface of elements where old concrete is strengthened with a new concrete overlay, there is no evidence from this study that steel connectors previously cast in the concrete are significantly more effective than those inserted afterwards. Equally the techniques used to increase the roughness of the substrate surface while fresh are not any more or less efficient than the methods used for hardened concrete, such as sandblasting, adopted in the present study. Finally, from this and from previous studies [1–5], the authors propose the following design procedure to be adopted in codes: (1) the roughness of the substrate surface should be quantified by means of a roughness parameter assessed with an equipment such as the laser roughness analyzer [3], instead of being qualitatively classified; (2) the cohesion and the friction coefficient should be calculated based on this roughness parameter, instead of being linked to the finishing treatment of the surface; (3) the design value of the shear stress at the interface should be first compared with the shear strength of the interface provided by adhesion only; and (4) in the case shear strength by adhesion was proven not to be enough, steel connectors should be designed and, in this case, adhesion should not be considered.
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Acknowledgements The authors are grateful to Sika, Hilti, Betão Liz, Cimpor and Secil for their kind collaboration in this research project. References [1] Júlio ES, Branco F, Silva VD. Concrete-to-concrete bond strength. Influence of the roughness of the substrate surface. Constr Build Mater 2004;18(9):675–81. [2] Santos P, Júlio E, Silva VD. Correlation between concrete-to-concrete bond strength and the roughness of the substrate surface. Constr Build Mater 2007; 21(8):1688–95. [3] Santos P, Júlio E. Development of a laser roughness analyzer to predict in situ concrete-to-concrete bond strength. Mag Concr Res 2008;60(5):329–37. [4] Júlio ES, Branco F, Silva VD. Concrete-to-concrete bond strength, Influence of an epoxy-based bonding agent on a roughened substrate surface. Mag Concr Res 2005;57(8):463–8. [5] Júlio ES, Branco F, Silva VD, Lourenço JF. Influence of added concrete on concrete-to-concrete bond strength. Build Environ 2006;41(12):1934–9. [6] Birkeland PW, Birkeland HW. Connections in precast concrete construction. ACI J 1966;63(3):345–67. [7] Mattock AH, Hawkins NM. Shear transfer in reinforced concrete—recent research. PCI J 1972;17(2):55–75. Precast/Prestressed Concrete Institute. [8] Loov RE. Design of precast connections, seminar organized by Compa International Pte, Ltd, 8 p., Singapore, September 1978. [9] Walraven J, Frenay J, Pruijssers A. Influence of concrete strength and load history on the shear friction capacity of concrete members. PCI J 1987;32(1): 66–84. Precast/Prestressed Concrete Institute. [10] Randl N. Investigations on transfer of forces between old and new concrete at different joint roughness, Ph.D. thesis. Austria: University of Innsbruck; 1997. 379 p [in German].
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