Advertisement

Bulletin of Earthquake Engineering

, Volume 17, Issue 12, pp 6645–6670 | Cite as

Experimental investigation on the deformation capacity of lap splices under cyclic loading

  • Danilo TarquiniEmail author
  • João Pacheco de Almeida
  • Katrin Beyer
S.I.: Nonlinear Modelling of Reinforced Concrete Structural Walls

Abstract

Correct detailing and positioning of lap splices is essential in order to prevent premature failure of reinforced concrete structural members. Especially before the introduction of capacity design guidelines, lap splices were often placed in member regions that undergo inelastic deformations under earthquake loading. When assessing the seismic performance of such members, not only the lap splice strength, which was assessed in previous studies, but also information on the deformation capacity of lap splices is required. This paper analyses the results of a recently concluded experimental programme on spliced RC wall boundary elements tested under uniaxial tension–compression cyclic loading. The study aimed at investigating the influence of lap splice length, confining reinforcement and loading history on the deformation capacity of lap splices. The latter is defined as the average strain, at the onset of splice failure, ascribed to deformations originating from the lap splice zone. Analysis of the test results showed that the deformation capacity of lap splices: (1) increases with lap splice length; (2) increases with confining reinforcement but the effectiveness of the confining reinforcement is dependent on the lap splice length; (3) decreases with larger imposed compression levels; (4) is larger for bottom-casted with respect to top-casted lap splices. Finally, an empirical model is proposed to estimate the strain capacity of lap splices, which provides a good fit with the experimental results.

Keywords

Uniaxial cyclic tension–compression tests Lap splices Deformation capacity Lap splice length Confining reinforcement Loading history 

1 Introduction

Splicing of longitudinal reinforcement is unavoidable in reinforced concrete (RC) structures and it can be found in all types of structural members such as beams, columns and walls. If not appropriately detailed and/or located in regions where inelastic deformations are expected, the presence of the lap splices may lead to a significant reduction of the strength and/or displacement capacity of the structural member. As reviewed in the following section, past research mainly focused on the evaluation of the strength capacity of lap splices, mostly through monotonic tests on beams spliced in the constant moment region. Tests on RC columns or walls are scarcer as well as tests performed under cyclic loading. To the authors’ knowledge, no experimental programme aimed primarily at the characterization of the deformation capacity of lap splices; and so despite its relevance in the context of performance-based design and assessment of structures, where displacement capacity rather than forces are compared with the seismic demand. The European seismic code (EN1998-3 2005) accounts for the presence of inadequately detailed lap splices in potential plastic hinge regions of RC members by reducing their ultimate chord rotation. However, due to the lack of specific experimental data, the proposed equation is obtained through a semi empirical approach (Biskinis and Fardis 2010). Similarly, other expressions available in the literature to compute the deformation capacity of lap splices are derived from extremely limited test databases (Hannewald 2013; Priestley et al. 1996). To address this gap, the present paper outlines the result of a recently concluded experimental programme focusing on the deformation capacity of lap splices. 24 test units, of which 22 with lap splices, designed to represent the boundary elements of spliced RC walls, were tested under uniaxial tension–compression cyclic loading. The specimens differed in terms of lap splice length, confining reinforcement, and loading history, as these are among the parameters most influencing the deformation capacity of lap splices, as identified in preceding studies (Almeida et al. 2017b; Tarquini et al. 2017). All specimens were instrumented to continuously monitor several displacement quantities as well as the applied forces. In particular, a dense mesh of optical sensors allowed to identify and isolate the lap splice deformation contribution from the total imposed displacement.

The present document is organized as follows. Past experimental tests on RC members with lap splices are described in Sect. 2, which is then complemented with the summary of the authors’ recently concluded experimental programme in Sect. 3. Section 4 discusses the influence of the main investigated experimental parameters on the deformation capacity of lap splices. In Sect. 5, based on experimental data, an empirical model for predicting the strain at failure of lap splices is proposed and validated. Conclusions are drawn in Sect. 6.

2 Experimental tests on RC members with lap splices: literature review

Early investigations on spliced RC members were conducted on beams with lap splices in the constant moment region, loaded monotonically up to failure (Chamberlin 1958; Chinn et al. 1955; Ferguson and Breen 1965; Tepfers 1973). The objective was to improve the understanding of the observed splitting failure modes as well as to quantify the strength of lap splices. The presence and amount of confining reinforcement, longitudinal rebar diameter, spacing and length of the splices, concrete cover and concrete strength were considered among the analysed parameters. Additional experimental work on the strength of lap splices was done by Ferguson and Briceno (1969), Ferguson and Krishnaswamy (1971), and Thompson et al. (1975), who used constant section beams to simulate spliced retaining walls: longitudinal rebars of large diameters, staggered splices, and splices in both regions of constant and non-constant moment were tested. Based on a regression analysis of the results obtained from a selection of the above mentioned beam monotonic tests, Orangun et al. (1977) proposed an equation for the bond strength of lap splices, which is to date a major reference for estimating this quantity.

The first experimental study on the cyclic inelastic behaviour of lap splices was carried out at Cornell University (Gergely and White 1980; Lukose et al. 1982) at the beginning of the 80's. Cyclic tests were motivated by the fact that, due to the scarcity of information on the performance of lap splices subjected to cyclic loading, most seismic codes of that time did not allow such details at locations of inelastic deformation, or else specified highly conservative design procedures. 68 large beams and 24 columns were tested with varying confining reinforcement and loading history (repeated and reversed cyclic loading). It was observed that adequate confining reinforcement was more effective for cyclic rather than monotonic loading, and that reversed cyclic loading led to earlier splice failure than repeated loading. The influence of a moment gradient along the lap splice was also investigated and its beneficial effect on the splices performance was recognized. Further experimental programmes on spliced RC beams were performed by Rezansoff et al. (1988, 1991, 1993), Sakurada et al. (1993), and Sparling and Rezansoff (1986), who studied the influence of confining reinforcement on the cyclic bond strength of lap splices; it was observed that adequately confined lap splices were sufficiently ductile to withstand a limited number of load reversals. Aristizabal-Ochoa (1983) carried out inelastic cyclic tests on spliced RC columns under reversed axial loads and observed a reduction in the attained maximum ductility with the number of cycles. The author also noticed that proper placement of transverse reinforcement is crucial in order to avoid brittle failure. Paulay (1982) proposed a simple design procedure to ensure that splices can sustain several cycles in the inelastic range; 8 RC columns subjected to lateral static loading were tested for validation purposes. Specific investigation on the behaviour of non-contact tension splices was carried out by Sagan et al. (1988) and Hamad and Mansour (1997), who conducted tests on 47 flat plate and 17 slab specimens, respectively. They concluded that the splice bar spacing influences the ultimate load carrying capacity of the structural member as well as the number of resisted inelastic tensile load cycles.

An equation for the strength of lap splices in compression based on monotonic axial column tests by Cairns and Arthur (1979) and Pfister and Mattock (1963) was proposed for the first time by Cairns (1985). It was claimed that the strength of tension and compression lap splices is influenced by the same factors (spliced length, confining reinforcement, longitudinal rebar diameter and concrete cover), although their relative importance is different. Cyclic tests on three columns and four beams under repeated compression loading were carried out by Panahshahi et al. (1992), who concluded that compression lap splices can be designed to sustain several cycles of inelastic loading. More recently, Chun et al. (2010, 2011) investigated the monotonic behaviour of compression splices in normal and high strength concrete while Askar (2016) studied the influence of splice length, transverse reinforcement, and end bearing conditions on spliced RC columns loaded monotonically up to failure.

In the last three decades, following major earthquakes in California (e.g. San Fernando 1971 or Loma Prieta 1989) that emphasized the vulnerability of spliced RC piers and columns, several experimental programmes were carried out focusing on strengthening techniques of such structural members. Chai et al. (1991) and Aboutaha et al. (1996) evaluated the effectiveness of steel jackets to improve the strength and ductility of piers and columns with short lap splices. Several retrofitting techniques such as welding of spliced bars, confining the splice region with steel angles or providing additional reinforcing bar ties were investigated by Valluvan et al. (1993), while seismic retrofit using prefabricated composite jacketing was carried out by Xiao and Ma (1997). Cyclic tests on non-retrofitted RC columns with details typical of pre mid-1970 US construction practice and subjected to increasing lateral load can be found in the works by Lynn et al. (1996) and Melek et al. (2003). Strengthening of non-ductile RC columns with carbon fiber reinforced polymers (CFRP) was explored by several authors, e.g. (Harries et al. 2006; Juntanalikit et al. 2016), and compared with the use of additional internal steel ties (Elsouri and Harajli 2011).

A large experimental program comprising 83 spliced beam specimens aimed at determining the influence on bond strength of relative rib area and bar diameters was carried out by Darwin et al. (1996). Conventional and experimental deformation patterns for the rebars were evaluated. The continuation of the test series, with additional 64 beam tests, is reported in Zuo and Darwin (2000), where the obtained data are used to determine an empirical relation for the bond strength of lap splices. The latter is expressed as function of concrete strength, relative rib area, bar size, and confinement provided by both concrete and transverse reinforcement. A modification to the above equation was proposed by Ichinose et al. (2004) in order to better capture the size-effect in the bond strength observed in their own experimental programme.

Epoxy coated reinforcing bars are used whenever corrosion protection represents a principal design requirement for RC members. In the 1990s, a multitude of studies on the splice strength of epoxy coated reinforcing bars were carried out on beam monotonic tests (Choi et al. 1992; Cleary and Ramirez 1992; DeVries et al. 1991; Hester et al. 1993; Treece and Jirsa 1990). It was concluded that epoxy coating significantly reduces the bond strength of lap splices and that, if used, a modification factor for the splice length should be adopted. Starting from the mid-1990s, research effort was also put in investigating the bond strength of lap splices in high-strength concrete (HSC) (Azizinamini et al. 1995; El-azab and Mohamed 2014; Hamad and Itani 1999; Hamad and Machaka 1999; Hwang et al. 1995; Kadoriku 1994) and fibre reinforced concrete (Hamad et al. 2001; Harajli and Salloukh 1998; Lagier et al. 2015) beams. Finally, as smooth bars are regularly encountered in historical structures, several test series on RC members featuring spliced plain bars can also be found in the literature (e.g. Goksu et al. 2014; Hassan and Feldman 2012; Verderame et al. 2008a, b).

Tests on RC wall elements with lap splices are scarce and relatively recent if compared to those on columns and beams (Aaleti et al. 2013; Almeida et al. 2017a; Bimschas 2010; Elnady 2008; Hannewald et al. 2013; Layssi and Mitchell 2012; Lowes et al. 2012; Paterson and Mitchell 2003; Villalobos et al. 2017). The units were typically subjected to a constant axial load and increasing cyclic lateral displacements. A detailed review of the corresponding spliced RC walls, including observations on the main failure modes, can be found in Almeida et al. (2017b).

3 New experimental programme on RC members with lap splices

3.1 Test setup, units, and loading

A total of 24 RC members, 22 of which with lap splices and two reference units with continuous reinforcement, were tested under uniaxial tension–compression cyclic loading at the structural laboratory of the École Polytechnique Fédérale de Lausanne (EPFL). The experimental programme, as well as some relevant experimental observations, are herein summarized. A detailed description of the tests is available in a separate data paper (Tarquini et al. 2019), which also outlines the organization of the available experimental data, free to download from the Zenodo platform at the  https://doi.org/10.5281/zenodo.1205887.

The test units (TUs) represent spliced RC wall boundary elements and were designed based on the tests carried out by Bimschas (2010) and Hannewald et al. (2013). The geometry was common to all specimens and consisted of a column height h = 1260 mm and a square 200 × 200 mm cross section, as illustrated in Fig. 1. Foundation and top beam blocks of dimensions 550 × 550 × 300 mm were included to connect the TUs to the uniaxial testing machine (± 125 mm stroke and + 2.5/− 10 MN force capacity) via four T-shaped steel profiles. The longitudinal reinforcement was formed by four 14 mm diameter rebars (corresponding to a longitudinal reinforcement ratio Al ≈ 1.5%) which, except for the units with continuous reinforcement, were spliced above the foundation. Four different lap splice lengths were considered, spanning from 25 to 60 times the longitudinal bar diameter (Øl). The confining reinforcement was composed of 6 mm diameter stirrups with non-staggered 90° hooks (hook location for each TU is part of the available material for download, see Tarquini et al. (2019)) and were wired to the longitudinal reinforcement at the 4 corners. A hook length of 100 mm, corresponding to 16 times the stirrup diameter, was provided, as depicted in Fig. 1d. Confining reinforcement ratios ranged from 0 to 0.3%. Such intervals were chosen as representative of both pre-seismic and code-compliant central European construction practice. The longitudinal reinforcement was made of hot-rolled steel (fy = 510 MPa, εsh = 0.95‰, fu = 635 MPa, εu = 9.3‰) while cold formed steel was used for the transverse reinforcement (fy = 475 MPa, fu = 625 MPa, εu = 9.8‰). The TUs were casted horizontally four at a time and the 28-day concrete compressive strength, obtained from cylinder tests, remained between 30 and 35 MPa.
Fig. 1

Test setup, geometry and reinforcement layout of the unit LAP-P3 (dimensions in mm): a Photo of the TU before the test; b reinforcement content in the N-S direction; c reinforcement content in the E-W direction; d cross section

Five different uniaxial loading protocols were considered, one monotonic (M) and four cyclic (C1–C4). The cyclic histories featured increasing applied tension levels, with two cycles performed at each displacement amplitude. The attained compression levels differentiated the different protocols: a 10:1 ratio between the imposed tension and compression displacements was used in the main loading protocol, labelled C1. A 10:2 ratio characterized protocol C2, whereas C3 corresponded to a repeated cyclic loading history, i.e. the specimen was brought back to zero displacement after each applied tensile displacement amplitude. Finally, in protocol C4 an approximate axial load ratio (ALR) of 90% was applied at each compression load step.

The complete matrix of the tests is listed in Table 1. Unit-specific parameters such as lap splice length, confining reinforcement, loading history, and location of the top-casted face are reported along with some important test results including the observed failure mode (discussed in Sect. 3.2), maximum forces attained both in tension and compression, and the displacement at specimen failure. The latter was defined as the displacement corresponding to 20% loss of force capacity, either in tension or in compression.
Table 1

Test matrix

TU

ls (mm)

ls (× Øl)

At (mm)

ρt (%)

LH

TC face

FM

Nmax (kN)

Nmin (kN)

Δf (mm)

LAP-P1

560

40

Ø6@100

~ 0.3

C1

E

M

362

− 1074

24

LAP-P2

560

40

Ø6@300

~ 0.1

C1

W

M

323

− 568

5.3

LAP-P3

560

40

Ø6@200

~ 0.15

C1

E

S–U

326

− 682

9

LAP-P4

350

25

Ø6@100

~ 0.3

C1

W

S–U

301

− 513

5

LAP-P5

840

60

Ø6@300

~ 0.1

C1

E

M

357

− 1018

18

LAP-P6

560

40

Ø6@200

~ 0.15

M

W

S–U

321

[–]

5.5

LAP-P7

560

40

Ø6@200

~ 0.15

C1

W

S–U

319

− 523

6

LAP-P8

560

40

Ø6@200

~ 0.15

M

W

S–U

340

[–]

13

LAP-P9

560

40

Ø6@200

~ 0.15

C2

W

S–U

311

− 766

6

LAP-P10

840

60

Ø6@100

~ 0.3

C1

E

C–C

359

− 1163

− 2.7

LAP-P11

350

25

Ø6@200

~ 0.15

C1

E

S–U

252

− 281

3.9

LAP-P12

350

25

Ø6@300

~ 0.1

C1

E

S-U

254

− 286

3.9

LAP-P13

840

60

Ø6@200

~ 0.15

C1

W

C–C

363

− 973

− 2.1

LAP-P14

840

60

[–]

0

C1

E

S–E

342

− 835

9

LAP-P15

560

40

[–]

0

C1

E

S–E

322

− 315

5.8

LAP-P16

560

40

Ø6@150

~ 0.2

C1

E

S–U

340

− 832

10

LAP-P17

560

40

Ø6@120

~ 0.25

C1

E

S–U

351

− 911

15

LAP-P18

700

60

Ø6@200

~ 0.15

C1

W

M

355

− 1091

17

LAP-P19

560

40

Ø6@120

~ 0.25

M

E

M

333

[–]

17

LAP-P20

560

40

Ø6@120

~ 0.25

C1

W

S–U

342

− 1011

12.5

LAP-P21

560

40

Ø6@120

~ 0.25

C3

E

M

358

− 563

20

LAP-P22

560

40

Ø6@120

~ 0.25

C4

E

S–U

318

− 1211

9.4

LAP-C1

[–]

[–]

Ø6@200

~ 0.15

C1

E

C–C

353

− 1304

− 4

LAP-C2

[–]

[–]

Ø6@100

~ 0.3

C1

E

C–C

366

− 1342

− 3.9

TU test unit; ls lap splice length; Øl longitudinal bar diameter; At confining reinforcement content; ρt confining reinforcement ratio; LH loading history type (see text description); TC face location of top casted face; FM failure mode (see text description); Nmax maximum applied (tensile) force; Nmin minimum applied (compression) force; Δf global displacement at specimen failure

Load cells were used to monitor the applied axial forces, whereas both hard-wired and optical instrumentation systems were employed to evaluate global and local displacements. Namely: (1) four linear variable differential transducers (LVDTs) were installed at the column corners by means of a plumb line system; (2) six LVDTs arranged in two chains were located on the East and West unit faces; and (3) light-emitting diodes (LEDs) were glued to the North and South column sides, both on the concrete and on the spliced steel bars (through holes prepared during casting), according to a regular mesh.

3.2 Experimental observations

The cracking behaviour of the TUs and the observed failure modes are briefly discussed in the following two subsections and framed into general categories. For a more detailed and unit-specific description of the TUs’ behaviour, as well as to consult their hysteretic response, the reader is referred to the data paper (Tarquini et al. 2019), which shares all experimental data in digital form.

3.2.1 Cracking behaviour

Up until the failure of one of the splices or the occurrence of severe damage in compression such as concrete spalling or rebar buckling, all the specimens with spliced longitudinal reinforcement behaved in a similar way. At the application of the first tensile loading (Δv = 1 mm), six to ten horizontal cracks opened along the unit height while two formed at the top beam and foundation interfaces. Due to the sudden stiffness change owing to the varying longitudinal reinforcement content, two of these cracks were always located at the lap splice extremities (the bottom one therefore coinciding with the foundation interface crack). The cracks were roughly equally spaced along the entire length of the member and had a constant width (w ≈ 0.1 mm). Yielding of the longitudinal reinforcement occurred at an imposed displacement of about Δy ≈ 4 mm. In the previous tensile cycles (i.e., at Δv = 2 mm and Δv = 3 mm), a distinct crack width evolution was observed between the spliced region and the zone above the splices. In particular, crack widths within the spliced region remained rather constant (w ≈ 0.1 mm) while the others increased with the imposed tensile displacement (up to w ≈ 0.3 mm for Δv = 3 mm). At Δv = 3 mm, the first vertical splitting cracks appeared at the splice loaded ends, leading to steel–concrete debonding action. During the following loading cycle to Δv = 6 mm, yielding of the longitudinal rebars took place, and the specimen behaviour became unit-specific. Specimens with short lap splices (ls = 25 × Øl) failed before reaching the yield force with vertical cracks opening along the entire lapped length and causing the almost total loss of the member force capacity. For the other specimens with longer lap splices, increasing the imposed tensile displacement beyond yielding produced a continuation of the widening of the horizontal cracks outside the lap splice region while crack opening remained approximately constant within the spliced zone (Fig. 2a, b). Vertical cracks extended from the lap ends towards the centre with failure occurring when the debonded length equalized the lap splice length. No extension of the vertical (tension-produced) splitting cracks was observed in compression until average compression strains in the order of εv ≈ − 1.5% (Δv = − 1.8 mm) were applied. At this stage vertical crushing cracks appeared, as well as minor concrete spalling. Whenever larger compression deformations were reached, extensive spalling developed above the spliced region followed by rebar buckling and eventually concrete core crushing.
Fig. 2

Crack distribution and width along the specimen height. LAP-P1: a crack distribution; b crack widths for three distinct levels of tensile displacement; LAP-C1: c crack distribution; d crack widths for three distinct levels of tensile displacement

Specimens with continuous reinforcement showed an approximately uniform crack distribution along their height, with horizontal cracks spaced about 150–200 mm (Fig. 2c). The crack width was rather constant, increasing with the imposed tensile displacement, as shown in Fig. 2d. Upon load reversals, crack closure with no damage was observed up to high imposed compression strain levels (εv ≈ − 2%). At this point, vertical cracks formed between the main horizontal cracks, preceding longitudinal rebar buckling and concrete crushing. The latter was the governing failure mode for these benchmark units (with continuous reinforcement), leading to a loss of almost 80% of their load carrying capacity. Rebar rupture was attained in the last pulling cycle, i.e. performed after specimen failure in compression.

3.2.2 Failure modes

The failure modes observed for each TU are listed in Table 1. Specimen failure, intended as a 20% loss of the maximum tension or compression force attained during each test (i.e., the TU capacity), was caused either by failure of one or more lap splices or by core concrete crushing. Two different lap splice failures could be distinguished: splitting–unzipping (S–U) and splitting–explosive (S–E). The former was characterized by vertical splitting cracks, which, originating at the lap loaded ends gradually extended along the entire splice length. The progression of the crack opening depended on the confining reinforcement, with lower confinement ratios leading to a faster crack development. Lap splice failure occurred when the vertical cracks from the lap ends joined about the centre of the splice (Fig. 3a), causing a strength loss of around 80% of the rebar yield strength. When little or no confinement reinforcement was provided (ρt < 0.1%), almost no vertical cracks could be observed before a sudden and loud lap splice failure occurred. No residual splice force was available for these failing splices and, upon load reversal, complete spalling of the cover concrete was typically observed (Fig. 3b). The term ‘mixed’ failure mode (M) is used to identify the few cases in which lap splices of the same unit failed according to different failure modes.
Fig. 3

Failure modes: a splitting-unzipping failure of the NE splice of LAP-P5; b splitting-explosive failure of the SE splice of LAP-P14; c concrete crushing of LAP-P10; d close-up of concrete crushing of LAP-P10; e close-up of rebar rupture of LAP-P21

Finally, core concrete crushing (C–C) occurred when lap splices were long and well confined (ls = 60Øl, ρt > 0.15%) and for the reference units with continuous reinforcement. It was always preceded by extensive concrete spalling, stirrups opening, and rebar buckling which, for spliced members, took place right above the lap splice region (Fig. 3c, d). Rebar rupture was never the primary failure mode for any of the TUs and it was always attained after specimen failure, during the last tensile pulling cycle (Fig. 3e).

4 Discussion of the experimental results

As discussed in the Introduction, the main objective of this paper is to assess the deformation capacity of lap splices as function of the test-series variables, i.e. lap splice length, confining reinforcement, and loading history. The quantity used to characterize the deformation capacity of lap splices as well as the method employed for its calculation from experimental measurements are described in the next Sect. 4.1 whilst the influence of each variable is separately assessed in Sects. 4.2.1, 4.2.2 and 4.2.3.

4.1 Definition of the average lap splice strain capacity

The average lap splice strain capacity (εls) is defined as the average deformation, at the onset of failure, owing exclusively to the lap splice region. Deformation contributions external to the lapped zone are not accounted for, such as the strain penetration of the rebar anchored into the foundation or the slip between the concrete and the rebar developed above the splices. In view of the available displacement measures in the spliced TUs (see Fig. 4b), the following steps were required to compute εls:
Fig. 4

Auxiliary plots and sketches for definition of lap splice strain capacity εls: a Identification of the four load steps at splice failure LSf for TU LAP-P15; b Location of optical markers around the lap splice region; c Strain distribution for the pair of spliced bars; d Validation of the assumption on the parameter α (α1 ≈ α2 ≈ 0.5)

  1. (a)

    Identify the load step at the onset of lap splice failure (LSf, see Fig. 4a);

     
  2. (b)
    Compute at LSf the lap splice displacement including the contribution of the two major cracks forming at the top and bottom lap splice ends (Δls,out). Referring to Fig. 4b, Δls,out is calculated as:
    $$\Delta_{ls,out} = \Delta_{F} - \Delta_{T}$$
    (1)
    where \(\Delta_{F}\) is the vertical displacement of foundation marker F and \(\Delta_{T}\) is the vertical displacement of top marker T. Note that positive displacement values are downwards (the piston is attached to the foundation) as the reference system in Fig. 4b indicates.
     
  3. (c)
    Calculate the width of the top and bottom end cracks (wTOP and wBOT, respectively):
    $$w_{TOP} = \Delta_{T'} - \Delta_{T}$$
    (2)
    $$w_{BOT} = \Delta_{F} - \Delta_{F'}$$
    (3)
     
  4. (d)
    Subtract the portion of wTOP and wBOT related to deformations occurring outside the lap splice region; namely, the part of the crack width due to the slip between the bar and the concrete above the lapped zone (\(\alpha_{1} \cdot w_{TOP}\)) and the part of the crack width caused by strain penetration in the foundation-anchored rebar (\(\alpha_{2} \cdot w_{BOT}\)). The processed lap splice displacement (Δproc) can therefore be expressed as:
    $$\Delta_{proc} = \Delta_{ls,out} - \alpha_{1} \Delta w_{TOP} - \alpha_{2} \Delta w_{BOT}$$
    (4)
     
  5. (e)
    Finally, the average lap splice strain capacity εls is the ratio between Δproc and the nominal lap splice length ls:
    $$\varepsilon_{ls} = \frac{{\Delta_{proc} }}{{l_{s} }}$$
    (5)
     
The calibration of the parameters α1 and α2 was carried out by comparing the displacement Δls,out and the splice-end-crack widths wTOP and wBOT with the displacement Δint obtained by integrating, along the lap splice length, the envelope of the top and bottom anchored rebar strains (Fig. 4c). The parameter α (assuming α1 = α2) was computed as follows:
$$\alpha= (\Delta_{ls,out} - \Delta_{int} )/\left( {w_{TOP} + w_{BOT} } \right)$$
(6)

The comparison is represented in Fig. 4d, which shows that α1 = α2 = 0.5 provided a good assumption for the present series of TUs. Further details on the calibration procedure as well as a mechanical interpretation of the obtained values for α1 and α2 are provided in the “Appendix”.

4.2 Influence of variable test parameters on the average strain capacity of lap splices

4.2.1 Lap splice length

Three groups of TUs were selected in order to investigate the influence of splice length (ls) on the deformation capacity of lap splices (εls). The units within each group varied only with regard to ls and featured the same confining reinforcement ratio (ρt) and imposed loading history (LH). The latter was common to all three groups and corresponded to C1.

The force–displacement responses of the selected TUs are displayed in Fig. 5a. The total applied axial force N is given on the vertical y-axis while the global column vertical displacement Δv and strain εv are reported on the bottom and top x-axes. A separate plot is provided for each combination of ls and ρt, where units belonging to the same group are represented with the same line colour. Namely blue, green and red are used for reinforcement ratios ρt = 0.1%, ρt = 0.15% and ρt = 0.3% respectively. Results of longer and more confined lap splices can be found in the downward and rightward directions. Separate markers are used to indicate the onset of each lap splice failure, with distinct shape and fill depending on the observed failure type. A black dashed line identifies the overall specimen failure (corresponding to a loss of around 20% of the specimen load capacity) whereas rebar rupture, which always occurred after specimen failure, is displayed with a black cross. A black circle specifies lap splices considered as outliers, which are characterized by a displacement capacity far larger than the one observed in all other lap splices of the same TU. This larger ductility of specific lap splices is related to the increased confining action provided by the stirrups, which, after failure of the first three splices, start to tilt with respect to their original horizontal position (see Fig. 3e). Such inclination introduces a compressive axial force component in the confining reinforcement contributing to restraining the pair of spliced rebars. However, since this effect is unit-specific and may not occur for different cross sectional geometries or reinforcement layouts, such outliers will not be considered in the analysis of the deformation capacity of lap splices.
Fig. 5

Influence of ls on the behaviour of lap splices: (a) Force–displacement responses; (b), (c), (d) Lap splice strain capacity for fixed confining reinforcement (ρt = 0.1%, ρt = 0.15%, ρt = 0.3%); (e) Lap splice strain capacity as function of ls (TC: top-casted rebar, BC: bottom-casted rebar)

The hysteretic curves of Fig. 5a show that the displacement capacity of spliced members increases with ls (i.e., the black dashed line consistently moves to larger displacements for longer splices). Except for ls = 60Øl and ρt > 0.15%, where the lap splices did not fail but concrete crushing was the predominant failure mode, specimen failure was always caused by the failure of at least one lap splice. For short splices (ls = 25Øl), simultaneous failure of the four lap splices took place before reaching the yielding of the longitudinal rebars. In all other cases, lap splice failure was observed after yielding of the longitudinal reinforcement and typically two splices failed clearly before the remaining two. The casting position is responsible for this difference as the increased porosity and water content reduces the concrete strength in top-casted splices (Luke et al. 1981); this issue will be addressed more in detail in the following paragraph. The responses of the two reference units with continuous reinforcement are plotted in grey. It can be observed that, although specimens featuring long and well confined splices (ls = 60Øl and ρt > 0.15%) did not fail, they nonetheless showed a significantly reduced deformation capacity with respect to the reference units. This is explained by the fact that, in spliced members, the deformation concentrates outside the lap splice region while cracks remain very small within the lapped zone.

The strain capacity of the single lap splices is plotted separately, for the three groups of TUs, in figures Fig. 5b, c and d. Outliers (namely, the last-holding splices of LAP-P1 and LAP-P3) and splices failing after crushing of the specimen were discarded. An additional marker (“+” symbol) is used to identify top-casted splices. The three plots show that εls increases almost linearly with ls, regardless of the provided confining reinforcement. Moreover, it is possible to distinguish the detrimental effect of top casting on the lap splice performance. In fact, except for short splices (ls = 25Øl) where the four failed simultaneously, lap splices located on the top casted face showed considerably reduced εls compared to those located on the bottom-casted face. However, it is important to point out that the determination of εls is less reliable for bottom-casted than top-casted splices as the specimen is in a more damaged state. In fact, after the failure of the first splice, the bottom and top crack width begin to differ and no redundant measure is available to double check the computed value of εlsint is not available, as discussed in Sect. 4.1).

Strains εls for all the selected TUs are showed altogether in Fig. 5e under the form of error-bar plot. Strains belonging to top-casted (TC) and bottom-casted (BC) splices are again differentiated. The shaded regions represent the scatter for strains corresponding to a certain combination of ls, ρt, and casting position. Again it is possible to observe the direct and almost-linear proportionality between εls and ls. Also, it is apparent that the scatter in the experimentally-determined strains increases with ls and when passing from top- to bottom-casted splices (i.e. with the imposed displacement demand and therefore with the damage state of the specimen). Please note that only one value of εls is available for the bottom-casted splices of specimen LAP-P1 (ls = 40Øl and ρt = 0.3%), which explains the zero data scatter.

4.2.2 Confining reinforcement

The influence of the confining reinforcement on the behaviour of lap splices is illustrated in Fig. 6, which contains similar plots to Fig. 5. The selected TUs were subjected to the same loading protocol (C1) and subdivided in three groups, each one referring to a fixed value of ls. Blue, green, and red line colours identify ls = 25Øl, ls = 40Øl, and ls = 60Øl respectively.
Fig. 6

Influence of ρt on the behaviour of lap splices: a force–displacement responses; b, c, d lap splice strain capacity for fixed lap splice length (ls = 25Øl, ls = 40Øl, ls = 60Øl); e lap splice strain capacity as function of ρt

Figure 6a shows that an increase in ρt does not necessarily correspond to a larger specimen ductility, especially for short lap splice lengths. In fact, the three TUs with ls = 25Øl depicted very similar hysteretic response and displacement at failure, although ρt varied from 0.1 to 0.3%. On the other hand, for long lap splices (ls = 60Øl), a direct proportionality between ρt and the displacement capacity of the TUs can be observed. It is noted that for specimens LAP-P13 and LAP-P10, the provided ρt was sufficient to promote a change in failure mode, from lap splice failure to concrete crushing. As for medium length lap splices (ls = 40Øl), an increase in ρt resulted in larger member ductility only for ρt > 0.15%. In fact, all units with ρt < 0.15% (LAP-P15, LAP-P2, LAP-P3 and LAP-P7) showed similar hysteretic response and displacement capacity; on the contrary, the latter increased for larger values of ρt in units with ρt > 0.15% (LAP-P16, LAP-P17, LAP-P20 and LAP-P1). Similar considerations to those above apply to the relation between ρt and εls, as displayed in figures Fig. 6b, c, and d: εls is approximately constant in Fig. 6b (note the scale of the vertical axis), monotonically increasing in Fig. 6d, and approximately bilinear (constant for ρt < 0.15%, and increasing for ρt > 0.15%) in Fig. 6c. Again, except for the case of short lap lengths (ls = 25Øl), within the same unit the top-casted splices always failed before the bottom-casted. An overall picture of the relationship between ρt and εls is given in Fig. 6e, which also confirms that data dispersion increases with the imposed displacement demand.

4.2.3 Loading history

Two set of TUs having as only variable parameter the imposed loading history (LH) were characterized by: ls = 40Øl and ρt = 0.15%—depicted with a blue colour line in Fig. 7—and ls = 40Øl and ρt = 0.25%—identified by a green colour line. Each group was composed by five TUs and included the reference cyclic (C1) and the monotonic (M) loading protocols. Tests on units with same configuration and applied loading histories (LAP-P6/P8, LAP-P3/P7, and LAP-P17/P20) were performed in order to assess the repeatability of the test results.
Fig. 7

Influence of LH on the behaviour of lap splices: a Force–displacement responses; b, c lap splice strain capacity for fixed lap splice length and confining reinforcement (ls = 40Øl and ρt = 0.15%, ls = 40Øl and ρt = 0.15%); d lap splice strain capacity as function of LH

Figure 7a shows a comparable displacement capacity of the TUs of the first set (ls = 40Øl and ρt = 0.15%), indicating that the imposed loading protocol did not play a significant role. Failure was in all cases triggered by splitting unzipping of the lap splices, occurring relatively soon after yielding of the longitudinal rebars. The smallest ductility is displayed by specimen LAP-P9, subjected to a loading protocol (C2) in which the compression levels were doubled with respect to those of C1. The two monotonic tests showed a significant aleatory difference in the attained displacement capacity, one failing right after yielding of the longitudinal reinforcement (LAP-P6, Δv ≈ 6 mm) and the other reaching Δv ≈ 12 mm. This scatter reduces slightly when εls is considered—see Fig. 7b, d, as the largest part of the member deformation occurs outside the lap splice region; recall that the crack width in the lapped zone remain relatively constant and small. Finally, the displacement and lap splice strain capacities of the two specimens tested under LH = C1 (i.e., LAP-P3 and LAP-P7) are in between the bounds obtained from the monotonic tests.

The impact of LH on the displacement capacity of the TUs is more clear for the second group (ls = 40Øl and ρt = 0.25%). It stands out that, in a cyclic loading, larger imposed compression levels reduce the deformation capacity of lap splices (Fig. 7a). In fact, the displacement at failure of LAP-P21 (LH = C3, repeated cyclic loading) is larger than that of LAP-P17 and LAP-P20 (LH = C1), which in turn are larger than the one of LAP-P22 (LH = C4, fixed compression level N ≈ 0.9 ALR). Comparing lap splices according to their order of failure, a similar trend applies to εls, as displayed in Fig. 7c and d. For this configuration of TUs, a single test was performed under monotonic loading (LAP-P19), which attained a strain capacity in between those of tests with protocols C1 and C3.

While the available test results allow such qualitative comments on the influence of the loading history on the deformation capacity of lap splices, further testing is required for appropriate quantitative characterization. Moreover, it was observed that, among the considered parameters, LH was the one with lowest impact on the lap splice strain capacity. Therefore, in the expression to estimate εls proposed in the next section, the effect of LH will be disregarded.

5 Prediction of the lap splice strain capacity

Based on the experimental results discussed above, a predictive equation for εls is proposed in this section, which builds on the following observations: (1) εls increases with ls, regardless of the provided ρt; (2) the influence of confining reinforcement on εls depends on ls. In particular, for short lap splices (ls = 25Øl), εls is not affected by variations in ρt; on the other hand, for long lap splices (ls = 60Øl), a small change from ρt = 0 to ρt ≈ 0.1% produces a significant increase of εls; finally, for medium lap lengths (ls = 40Øl) an increase of εls is observed only for provided ρt > 0.15%; (3) loading history is the least influential among the considered variable parameters and further testing is required for an appropriate quantification; (4) casting position plays a significant role in the displacement capacity of lap splices, with lower εls associated to bottom-casted splices.

The expression for εls is defined in the two-variable space ls and ρt, with 25Øl < ls < 60Øl and 0 < ρt < 0.3%. Stemming from the above consideration (2), two regions were defined in this [ls, ρt] domain: one in which an increase in ρt causes an increase in εls (labelled subdomain A) and another where the strain capacity does not depend on the confining reinforcement (subdomain B). The separation between the two regions is specified by the following equation:
$$\frac{{l_{s} }}{{\emptyset_{l} }} + \frac{60 - 25}{0.3} \cdot \rho_{t} - 60 = 0$$
(7)
where ρt is defined in percentage (%). The previous expression represents the line passing through the points with coordinates [ls, ρt] = [60Øl, 0%] and [25Øl, 0.3%]. All combinations of [ls, ρt] leading to positive values for Eq. (7) fall into subdomain A while those resulting in negative values belong to subdomain B. Moreover, since for ls = 60Øl and ρt > 0.15% no lap splice failure was observed, the equation should not be used in that subrange. A linear equation in the two variables ls and ρt is then fitted, for each subdomain, through the experimental εls values. The following two systems of equations, for top- and bottom-casted lap splices, were respectively obtained:
$$\left\{\!\!\begin{array}{*{20}l}\varepsilon_{ls}^{TC} = - 23 + 50 \cdot \rho_{t} + 0.44 \cdot \frac{{l_{s} }}{{\emptyset_{l} }} &\quad \to \, \left[ {l_{s} ,\rho_{t} } \right] \,\in \hbox{subdomain}\, \hbox{A}\\ \varepsilon_{ls}^{TC} = 1.2 + 0.04 \cdot \frac{{l_{s} }}{{\emptyset_{l} }} & \quad\to \, \left[ {l_{s} ,\rho_{t} } \right] \,\in \hbox{subdomain}\, \hbox{B} \end{array}\right.$$
(8, 9)
$$\left\{\!\!\begin{array}{*{20}l}\varepsilon_{ls}^{BC} = - 36 + 70 \cdot \rho_{t} + 0.76 \cdot \frac{{l_{s} }}{{\emptyset_{l} }} & \quad\to \, \left[ {l_{s} ,\rho_{t} } \right] \,\in \hbox{subdomain}\, \hbox{A}\\\varepsilon_{ls}^{BC} = - 2.1 + 0.2 \cdot \frac{{l_{s} }}{{\emptyset_{l} }} & \quad\to \, \left[ {l_{s} ,\rho_{t} } \right]\, \in \hbox{subdomain}\, \hbox{B}\end{array}\right.$$
(10, 11)
where εls is expressed in permille (‰). From a geometrical viewpoint, Eqs. (8)–(11) represent two planes in the [ls, ρt] space. The two planes of each system intersect on a line whose projection on the plane εls = 0 is given by Eq. (7). The fit between the experimental and predicted values of εls is displayed, for top- and bottom-casted splices, in Fig. 8a and b as 3D plots. Residual plots are instead given in Fig. 8c and d. The predictive equation captures rather well the trend of the experimental data as well as the coupling between ls and ρt. The effect of the different loading protocols shows up in the relatively large scatter observable at combinations [ls = 40Øl; ρt = 0.15%] and [ls = 40Øl; ρt = 0.25%]. All values of experimental (εlsexp) and predicted (εlspred) lap splices strain capacities are listed in Table 2 together with the associated model error υ, defined as:
Fig. 8

Predictive equation for εls versus experimental results. Spatial representation for: a top-casted and b bottom-casted lap splices; residual plot for: c top-casted and d bottom-casted lap splices

Table 2

Experimental and predicted values of lap splice strain capacity, εls

TU

Top-casted splices

Bottom-casted splices

Locationa

Δproc (mm)

εlsexp (‰)

εlspred (‰)

Error (%)

Location

Δproc (mm)

εlsexp (‰)

εlspred (‰)

error (%)

LAP-P1

SE

4.68

8.36

8.74

4.5

NW

8.99

16.05

14.19

11.6

NE

5.65

10.09

13.4

SWb

n/a

n/a

n/a

LAP-P2

NW

1.48

2.64

2.80

5.9

NE

4.19

7.49

5.90

21.2

SW

1.37

2.45

14.5

SE

3.64

6.50

9.2

LAP-P3

NE

1.75

3.12

2.80

10.1

NW

4.22

7.54

5.90

21.7

SE

2.52

4.50

37.8

SWb

n/a

n/a

n/a

LAP-P4

NW

1.07

3.06

2.20

28.0

NE

1.06

3.04

2.90

4.7

SW

1.11

3.16

30.3

SE

0.93

2.67

8.6

LAP-P5

NE

6.45

7.67

8.11

5.7

NW

10.2

12.14

16.20

33.4

SE

5.58

6.64

22.1

SW

13.5

16.07

0.8

LAP-P6

NW

1.35

2.41

2.80

16.1

NEb

n/a

n/a

5.90

n/a

SW

1.36

2.43

15.3

SE

3.62

6.46

8.7

LAP-P7

NW

1.55

2.72

2.80

2.8

NE

3.54

6.32

5.90

6.7

SW

1.30

2.31

21.1

SE

3.50

6.26

5.7

LAP-P8

NW

2.17

3.88

2.80

27.7

NEb

n/a

n/a

5.90

n/a

SW

2.31

4.13

32.1

SE

4.74

8.46

30.3

LAP-P9

NW

1.19

2.13

2.80

31.8

NE

3.13

5.60

5.90

5.4

SW

1.15

2.05

36.3

SE

3.06

5.46

8.0

LAP-P11

NE

0.68

1.93

2.20

14.1

NW

0.6

1.71

2.90

69.2

SE

0.62

1.76

25.2

SW

0.7

2.00

45.0

LAP-P12

NE

0.72

2.04

2.20

7.7

NW

0.71

2.03

2.90

43.0

SE

0.63

1.80

22.2

SW

0.64

1.84

57.4

LAP-P14

NE

6.44

7.67

3.60

53.0

NW

5.94

7.07

9.90

40.0

SE

2.76

3.29

9.6

SW

3.17

3.77

162.3

LAP-P15

NE

1.69

3.02

2.80

7.2

NW

4.2

7.50

5.90

21.3

SE

1.4

2.50

12.0

SW

3.45

6.17

4.4

LAP-P16

NE

2.99

5.33

4.02

24.5

NW

4.67

8.34

7.59

8.9

SE

2.1

3.75

7.3

SW

4.19

7.48

1.5

LAP-P17

NE

3.32

5.92

6.38

7.8

NW

5.22

9.33

10.89

16.8

SE

3.44

6.14

3.9

SW

4.88

8.71

25.0

LAP-P18

NW

4.73

6.76

6.07

10.2

NE

7.62

10.89

11.90

9.3

SW

4.07

5.81

4.4

SE

9.7

13.86

14.2

LAP-P19

NE

2.15

3.83

6.38

66.6

NW

6.25

11.16

10.89

2.4

SE

2.54

4.54

40.7

SWb

n/a

n/a

n/a

LAP-P20

NW

3.3

5.89

6.38

8.3

NE

6.76

12.08

10.89

9.8

SW

2.88

5.14

24.1

SEb

n/a

n/a

n/a

LAP-P21

NE

4.66

8.32

6.38

23.3

NW

9.3

16.61

10.89

34.4

SE

4.06

7.25

12.0

SWb

n/a

n/a

n/a

LAP-P22

NE

3.7

6.61

6.38

3.4

NW

3.5

6.25

10.89

74.3

SE

1.825

3.26

95.8

SW

4.13

7.38

47.5

 

Error Avg: 20.9%

Error Avg: 26.1%

aCorner of the specimen where the considered splice was located during the testing

bOutlier data (see explanation in text)

$$\nu = \left| {1 - \frac{{\varepsilon_{ls}^{pred} }}{{\varepsilon_{ls}^{exp} }}} \right|$$
(12)

LAP-P13 and LAP-P10 are not included in the table since no lap failure was observed. The good match between the predicted and experimental εls is confirmed by an average error of 20% and 26% for top- and bottom-casted splices, respectively. Furthermore, the same quantity drops to 13% and 20% if only TUs tested under the main loading protocol (C1) are considered.

6 Conclusions

Past experimental investigations on spliced RC members was mainly directed towards the characterization of their strength rather than their deformation capacity. The vast majority of tests included beams under monotonic loading with lap splices in the constant moment region, typically designed to fail before yielding of the longitudinal reinforcement. No experimental studies focusing on the deformation capacity of lap splices, particularly beyond yielding, are currently available in the literature. This represents a limitation to the application of performance-based design and assessment philosophies, where displacements rather than forces are compared with the seismic demand.

This study presents the results of an experimental programme on spliced RC members. The 24 test units represented the boundary elements of RC walls typical of both pre-seismic and code-compliant central European construction practice, and were tested under uniaxial tension–compression cyclic loading. The aim of the test series was to investigate the influence of lap splice length, confining reinforcement, and loading history on the deformation capacity of lap splices. A total of four lap splice lengths, from 25 to 60 times the longitudinal diameter, five confining reinforcement ratios, from 0 to 0.3%, and five loading protocols, 4 cyclic and one monotonic, were considered. Two reference test units with continuous reinforcement were also tested. Extensive instrumentation, including optical sensors glued directly on the spliced pairs or rebars, continuously monitored displacements of a comprehensive grid of points on the specimen faces.

The experimental data are used to derive an expression for predicting the deformation capacity of lap splices, which is defined as the average strain, at the onset of splice failure, owing exclusively to deformation contributions from the lapped region. Anchorage slip due to strain penetration or rebar slip from member locations outside the lap splice region were thus removed. With this approach, a strain limit for each of the four lap splices of each test unit was determined. The equation for the lap splice deformation capacity accounts for the confining reinforcement ratio, the lap splice length (as function of the bar diameter) and the position of the bar during casting. Larger deformation capacities were reached by bottom-casted splices with respect to top-casted, underlining the importance of concrete quality. Compared to these three parameters, the loading history was found to have only a minor influence and the database is insufficient to quantify it. However, it was clear that larger imposed compression levels lead to a decrease in the splice failure strain.

The deformation capacity of lap splices increases with the splice length, irrespectively of the provided confining reinforcement. The effectiveness of the confining reinforcement depended instead on the lap splice length. Namely, the deformation capacity of short lap splices (ls = 25Øl) is insensitive to the confining reinforcement ratio; on the other extreme, even very low levels of confining reinforcement are sufficient to increase the deformation capacity of long lap splices (ls = 60Øl); for intermediate lap-splice lengths the splice deformation capacity increases with confining reinforcement only beyond a certain ratio (ρt > 0.15%). To account for this observation, two sets of equations are derived: the first one is applicable to long, well confined splices (referred to as subdomain A) and the second one to shorter, less confined or unconfined lap splices (subdomain B). Coefficients are determined for top- and bottom-casted splices yielding an average error of about 20 and 26% for the former and the latter, respectively.

The predictive equation should not be applied outside of the bounds of the tested domain of lap splice lengths (25Øl ≤ ls ≤ 60Øl) and confining reinforcement ratios (0 ≤ ρt ≤ 0.3%). Moreover, for the cases of ls ≥ 60Øl and ρt > 0.15%, no lap splice failure was observed; concrete crushing took place above the lap splice region preceded by concrete spalling, stirrups opening and rebar buckling. This confirms that, in RC walls with well-detailed lap splices, a relocation of the plastic hinge above the lap splice region might be expected, with consequent reduction of the member shear span. Additionally, rebar rupture at the foundation interface did not occur, arguably due to the absence of moment gradient, the specifics of lap splice configurations, and the adopted ratio between tension and compression strains. Finally, the confining reinforcement ratios used within this study to compute the strain capacity of lap splices refer to stirrups with 90° hooks. The proposed equation might be conservative for the assessment of current seismic-code-compliant RC members with 135° hooks.

Notes

Acknowledgements

The financial support by the Swiss Federal Roads Office (FEDRO) to the Project Number AGB 2015/002, under which the present study is carried out, is acknowledged. Moreover, the authors would like to thank Tiago Nico Pereira and Maria Katsidoniotaki for the precious help in the laboratory.

References

  1. Aaleti S, Brueggen BL, Johnson B, French CE, Sritharan S (2013) Cyclic response of reinforced concrete walls with different anchorage details: experimental investigation. J Struct Eng 139(July):1181–1191CrossRefGoogle Scholar
  2. Aboutaha RS, Engelhardt MU, Jirsa JO, Kreger MF (1996) Retrofit of concrete columns with inadequate lap splices by the use of rectangular steel jackets. Earthq Spectra 12:693–714CrossRefGoogle Scholar
  3. Almeida J, Prodan O, Rosso A, Beyer K (2017a) Tests on thin reinforced concrete walls subjected to in-plane and out-of-plane cyclic loading. Earthq Spectra 33(1):323–345CrossRefGoogle Scholar
  4. Almeida JP, Prodan O, Tarquini D, Beyer K (2017b) Influence of lap-splices on the cyclic inelastic response of reinforced concrete walls. I: Database assembly, recent experimental data, and findings for model development. J Struct Eng. 143(12)Google Scholar
  5. Aristizabal-Ochoa JD (1983) Earthquake resistant tensile lap splices. J Struct Eng 109(4):843–858CrossRefGoogle Scholar
  6. Askar HS (2016) An experimental investigation on contact compression lap splice in circular columns. HBRC J 12(2):137–146CrossRefGoogle Scholar
  7. Azizinamini A, Chisala M, Ghosh SK (1995) Tension development length of reinforcing bars embedded in high-strength concrete. Eng Struct 17(7):512–522CrossRefGoogle Scholar
  8. Bimschas M (2010) Displacement based seismic assessment of existing bridges in regions of moderate seismicity. Ph.D. Thesis, ETH Zurich (18849)Google Scholar
  9. Biskinis D, Fardis MN (2010) Flexure-controlled ultimate deformations of members with continuous or lap-spliced bars. Struct Concr 11(2):93–108CrossRefGoogle Scholar
  10. Cairns J (1985) Strength of compression splices: a reevaluation of test data. ACI J 82(4):510–516Google Scholar
  11. Cairns J, Arthur PD (1979) Strength of lapped splices in reinforced concrete columns. ACI J Proc 76(2):277–296Google Scholar
  12. Chai YH, Priestley MJN, Seible F (1991) Seismic retrofit of circular bridge columns for enhanced flexural performance. ACI Struct J 88(88):572–584Google Scholar
  13. Chamberlin SJ (1958) Spacing of spliced bars in beams. ACI J Proc 54(2):689–697Google Scholar
  14. Chinn J, Ferguson PM, Thompson JN (1955) Lapped splices in reinforced concrete beams. ACI J 52(10):201–213Google Scholar
  15. Choi C, Hadje-Ghaffari H, Darwin D, Mccabe SL (1992) Bond of epoxy-coated reinforcement: bar parameters. ACI Mater J 88(2):207–217Google Scholar
  16. Chun S, Lee S, Oh B (2010) Compression lap splice in unconfined concrete of 40 and 60 MPa (5800 and 8700 psi) compressive strengths. ACI J Proc 107(2):170–178Google Scholar
  17. Chun S, Lee S, Oh B (2011) Compression splices in high-strength concrete of 100 MPa (14, 500 psi) and less. ACI Struct J 108(6):715–724Google Scholar
  18. Cleary DB, Ramirez JA (1992) Bond strength of epoxy-coated reinforcement. ACI Mater J 88(2):146–149Google Scholar
  19. Darwin D, Tholen ML, Idun EK, Zuo J (1996) Splice strength of high relative rib area reinforcing bars. ACI Struct J 93:95–107Google Scholar
  20. DeVries RA, Moehle JP, Hester W (1991) Lap splice of plain and epoxy-coated reinforcements: an experimental study considering concrete strength, casting position, and anti-bleeding additives. Report No. UCB/SEMM-91/02 Structural Engineering Mechanics and Materials, University of California, Berkeley, California, 86Google Scholar
  21. El-azab A, Mohamed HM (2014) Effect of tension lap splice on the behavior of high strength concrete (HSC) beams. HBRC J 10(3):287–297CrossRefGoogle Scholar
  22. Elnady EMM (2008) Seismic rehabilitation of RC structural walls. Ph.D. Thesis, McMaster UniversityGoogle Scholar
  23. Elsouri AM, Harajli MH (2011) Seismic repair and strengthening of lap splices in RC columns: carbon fiber—reinforced polymer versus steel confinement. J Compos Construct 15(5):721–731CrossRefGoogle Scholar
  24. EN1998-3 (2005) Eurocode 8: design of structures for earthquake resistance. Part 3: Assessment and retrofitting of buildings. Doc CEN/TC250/SC8/N306. European Committee for StandardizationGoogle Scholar
  25. Ferguson PM, Breen JE (1965) Lapped splices for high strength reinforcing bars. ACI J Proc 62(9):1063–1068Google Scholar
  26. Ferguson PM, Briceno EA (1969) Tensile lap splices Part I: retaining wall type, varying moment zone. The Texas Highway Department, Research report No. 113-2Google Scholar
  27. Ferguson PM, Krishnaswamy CN (1971) Tensile lap splices, Part II: Design recommendations for retaining wall, Splices and large bar splices. The Texas Highway Department in cooperation with the U.S. Department of Transportation, Report 113-2 Cont., Federal Highway Administration, Bureau of Public RoadsGoogle Scholar
  28. Gergely P, White R (1980) Seismic design of lapped splices in reinforced concrete. In: Proceedings of the 7th world conference on earthquake engineering, Istanbul, vol 4, pp 281–288Google Scholar
  29. Goksu C, Yilmaz H, Chowdhury SR, Orakcal K, Ilki A (2014) The effect of lap splice length on the cyclic lateral load behavior of RC members with low-strength concrete and plain bars. Adv Struct Eng 17:639–658.  https://doi.org/10.1260/1369-4332.17.5.639 CrossRefGoogle Scholar
  30. Hamad BS, Itani MS (1999) Bond strength of reinforcement in high-performance concrete : the role of silica fume, casting position, and superplasticizer dosage. ACI Mater J 95(5):499–511Google Scholar
  31. Hamad BS, Machaka MF (1999) Effect of transverse reinforcement on bond strength of reinforcing bars in silica fume concrete. Mater Struct 32:468–476CrossRefGoogle Scholar
  32. Hamad BS, Mansour M (1997) Bond strength of noncontact tension lap splices. ACI Struct J 93(3):1–11Google Scholar
  33. Hamad BS, Harajli MH, Jumaa G (2001) Effect of fiber reinforcement on bond strength of tension lap splices in high-strength concrete. ACI Struct J 98(5):638–647Google Scholar
  34. Hannewald P (2013) Seismic behavior of poorly detailed RC bridge piers. Ph.D. Thesis, École Polytechnique Fédérale de LausanneGoogle Scholar
  35. Hannewald P, Bimschas M, Dazio A (2013) Quasi-static cyclic tests on RC bridge piers with detailing deficiencies. Institut fur Baustatik und Konstruktion, Bericht Nr. 352, ETH Zurich, SwitzerlandGoogle Scholar
  36. Harajli MH, Salloukh KA (1998) Effect of fibers on development/splice strength of reinforcing bars in tension. ACI Mater J 94(4):317–424Google Scholar
  37. Harries KA, Ricles JR, Pessiki S, Sause R (2006) Seismic retrofit of lap splices in nonductile square columns using carbon fiber-reinforced jackets. ACI Struct J 103(6):874–884Google Scholar
  38. Hassan MN, Feldman LR (2012) Behavior of lap-spliced plain steel bars. ACI Struct J 109(109):235–244Google Scholar
  39. Hester CJ, Salamizavaregh S, Darwin D, McCabe SL (1993) Bond of epoxy-coated reinforcement: splices. ACI Struct J 90(1):89–102Google Scholar
  40. Hwang S, Lee Y, Lee C (1995) Effect of silica fume on the splice strength of deformed bars of high-performance concrete. ACI Struct J 91(3):294–302Google Scholar
  41. Ichinose T, Kanayama Y, Inoue Y, Bolander JE (2004) Size effect on bond strength of deformed bars. Constr Build Mater 18(7):549–558CrossRefGoogle Scholar
  42. Juntanalikit P, Jirawattanasomkul T, Pimanmas A (2016) Experimental and numerical study of strengthening non-ductile RC columns with and without lap splice by carbon fiber reinforced polymer (CFRP) jacketing. Eng Struct 125:400–418CrossRefGoogle Scholar
  43. Kadoriku (1994) Study on behavior of lap splices in high-strength reinforced concrete members. Doctorate Thesis, Kobe University, JapanGoogle Scholar
  44. Lagier F, Massicotte B, Charron J (2015) Bond strength of tension lap splice specimens in UHPFRC. Construct Build Mater 93:84–94CrossRefGoogle Scholar
  45. Layssi H, Mitchell D (2012) Experiments on seismic retrofit and repair of reinforced concrete shear walls. In: Proceedings of the 6th international conference on FRP composites in civil engineering-CICE, Rome, Italy, June 13–15, pp 1–8Google Scholar
  46. Lowes LN, Lehman DE, Birely AC, Kuchma DA, Marley KP, Hart CR (2012) Earthquake response of slender planar concrete walls with modern detailing. Eng Struct 43:31–47CrossRefGoogle Scholar
  47. Luke JJ, Hamad BS, Jirsa JO, Breen JE (1981) The influence of casting position on development and splice length of reinforcing bars. Research report 242-1, Center for Highway Research, The University of Texas at AustinGoogle Scholar
  48. Lukose K, Gergely P, White R (1982) Behavior of reinforced concrete lapped splices for inelastic cyclic loading. ACI J Proc 79(5):355–365Google Scholar
  49. Lynn AC, Moehle JP, Mahin SA, Holmes WT (1996) Seismic evaluation of existing reinforced concrete building columns. Earthq Spectra 12(4):715–739CrossRefGoogle Scholar
  50. Marti P, Alvarez M, Kaufmann W, Sigrist V (1998) Tension chord model for structural concrete. Struct Eng Int 8(4):287–298CrossRefGoogle Scholar
  51. Melek M, Wallace J, Conte J (2003) Experimental assessment of columns with short lap-splice subjected to cyclic loads. PEER Report 2003/04, Pacific Earthquake Engineering Research Center. University of California, Berkeley, CAGoogle Scholar
  52. Orangun CO, Jirsa JO, Breen JE (1977) A reevaluation of test data on development length and splices. ACI J 74(3):114–122Google Scholar
  53. Panahshahi N, White RN, Gergely P (1992) Reinforced concrete compression lap splices under inelastic cyclic loading. ACI Struct J 89(2):164–175Google Scholar
  54. Paterson J, Mitchell D (2003) Seismic retrofit of shear walls with headed bars and carbon fiber wrap. J Struct Eng 129:606–614CrossRefGoogle Scholar
  55. Paulay T (1982) Lapped splices in earthquake-resisting columns. ACI J 79(6):458–469Google Scholar
  56. Pfister JF, Mattock AH (1963) High strength bars as concrete reinforcement, part 5. Lapped splices in concentrically loaded columns. J PCA Res Dev Lab 5(2):27–40Google Scholar
  57. Priestley MJN, Seible F, Calvi GM (1996) Seismic design and retrofit of bridges. Wiley, New YorkCrossRefGoogle Scholar
  58. Rezansoff T, Zacaruk JA, Topping R (1988) Tensile lap splices in reinforced concrete beams under inelastic cyclic loading. ACI Struct J 85(1):46–52Google Scholar
  59. Rezansoff T, Konkankar US, Fu C (1991) Confinement limits for tension lap slices under static loading. Can J Civ Eng 19(3):447–453CrossRefGoogle Scholar
  60. Rezansoff T, Akanni A, Sparling B (1993) Tensile lap splices under static loading: a review of the proposed ACI 318 code provisions. ACI Struct J 90(4):374–384Google Scholar
  61. Sagan VE, Gergely P, White RN (1988) The behaviour and design of noncontact lap splices subjected to repeated inelastic tensile loading. Technical report NCEER-88-0033, Department of Structural Engineering, Cornell University, Ithaca, New YorkGoogle Scholar
  62. Sakurada T, Morohashi N, Tanaka R (1993) Effect of transverse reinforcement on bond splitting strength of lap splices. Trans Jpn Conc Inst 15:573–580Google Scholar
  63. Sezen H, Setzler EJ (2008) Reinforcement slip in reinforced concrete columns. ACI Struct J 105(3):280–289Google Scholar
  64. Sparling B, Rezansoff T (1986) The effect of confinement on lap splices in reversed cyclic loading. Can J Civ Eng 13(6):681–692CrossRefGoogle Scholar
  65. Tarquini D, Almeida JP, Beyer K (2017) Influence of lap-splices on the cyclic inelastic response of reinforced concrete walls. II: Shell element simulation and equivalent uniaxial model. ASCE J Struct Eng.  https://doi.org/10.1061/(ASCE)ST.1943-541X.0001859 CrossRefGoogle Scholar
  66. Tarquini D, Almeida JP, Beyer K (2019) Uniaxial cyclic tests on reinforced concrete members with lap splices. Earthq Spectra 35(2):1023–1043.  https://doi.org/10.1193/041418EQS091DP CrossRefGoogle Scholar
  67. Tastani SP, Brokalaki E, Pantazopoulou SJ, Asce M (2015) State of bond along lap splices. J Struct Eng 141(10):1–14CrossRefGoogle Scholar
  68. Tepfers R (1973) A theory of bond applied to overlapped tensile reinforcement splices for deformed bars. Calmers University of Technology, GoteborgGoogle Scholar
  69. Thompson MA, Jirsa JO, Breen JE, Meinheit DF (1975) The behavior of multiple lap splices in wide sections. Research report 154-1, Center for Highway Research, The University of Texas at AustinGoogle Scholar
  70. Treece RA, Jirsa J (1990) Bond strength of epoxy-coated reinforcing bars. ACI Mater J 86(2):167–174Google Scholar
  71. Valluvan R, Kreger ME, Jirsa JO (1993) Strengthening of column splices for seismic retrofit of nonductile reinforced concrete frames. ACI Struct J 90(4):432–440Google Scholar
  72. Verderame GM, Fabbrocino G, Manfredi G (2008a) Seismic response of R. C. columns with smooth reinforcement. Part I: monotonic tests. Eng Struct 30:2277–2288CrossRefGoogle Scholar
  73. Verderame GM, Fabbrocino G, Manfredi G (2008b) Seismic response of R. C. columns with smooth reinforcement. Part II: cyclic tests. Eng Struct 30:2289–2300CrossRefGoogle Scholar
  74. Villalobos E, Escolano-Margarit D, Ramìrez-Màrquez AL, Pujol S (2017) Seismic response of reinforced concrete walls with lap splices. Bull Earthq Eng 15(5):2079–2100CrossRefGoogle Scholar
  75. Xiao Y, Ma R (1997) Seismic retrofit of RC circular columns using prefabricated composite jacketing. J Struct Eng 123(10):1357–1364CrossRefGoogle Scholar
  76. Zuo J, Darwin D (2000) Splice strength of conventional and high relative rib area bars in normal and high-strength concrete. ACI Struct J 97(4):630–641Google Scholar

Copyright information

© Springer Nature B.V. 2019

Authors and Affiliations

  1. 1.Earthquake Engineering and Structural Dynamics Laboratory (EESD), School of Architecture, Civil and Environmental Engineering (ENAC)École Polytechnique Fédérale de Lausanne (EPFL)LausanneSwitzerland
  2. 2.Institute of Mechanics, Materials and Civil EngineeringUniversité catholique de LouvainLouvain-la-NeuveBelgium
  3. 3.Earthquake Engineering and Structural Dynamics Laboratory (EESD), School of Architecture, Civil and Environmental Engineering (ENAC)École Polytechnique Fédérale de Lausanne (EPFL)LausanneSwitzerland

Personalised recommendations