1. Introduction
During strong seismic events, the common regions where reinforced concrete beams and columns are connected are subjected to significantly higher stresses with respect to the stresses developed at the linked members [
1]. Therefore, if the beam–column joints are not appropriately designed to withstand these stresses, brittle failure of the joint regions can occur, resulting in a partial or even general collapse of the structure. This is particularly common in existing reinforced concrete (RC) structures, which were not designed according to the capacity design philosophy, and it is also a possibility even for the modern RC structures designed to satisfy the recommendations of Eurocode 2 and 8 [
2,
3] for ductility class medium (DCM) [
4].
When the anchorage of the beam longitudinal reinforcement in an exterior beam–column connection is artfully detailed and provides sufficient bonding between the bars and the surrounding concrete, significant horizontal shear stresses are placed on the joint region during cycling. For favorable steel–concrete bond conditions, the same is also true for the column longitudinal rebars passing through the joint, which introduce significant vertical shear stresses to the joint. Consequently, brittle shear failure of the connection will be triggered immediately after the shear stresses introduced to the joint exceed its shear capacity [
5,
6]. It is worth emphasizing that damage evolution and the extensive degradation of the overall bearing capacity are extremely rapid when the joint is totally unconfined or poorly confined [
7,
8]. Nevertheless, the joint shear failure cannot be effectively prevented solely by adding ties and by ensuring that the capacity design ratio value is higher than 1.30 (according to Eurocode 8), unless the developed shear stresses are limited to one half of the joint shear capacity (or lower) [
4,
9,
10]. Otherwise, the evolution of shear damage in the beam–column joint may be delayed but, eventually, the desirable elastic–plastic failure mechanism with plastic hinges formed solely in the beams will not be achieved.
The anchorage deficiency and degradation of bonding conditions significantly affect the seismic performance of the beam–column connection, since the shear strain to which a joint is subjected is directly related to the bond stresses developed between the beam and column rebars passing through the joint (or anchored inside it) and the surrounding concrete. For this reason, civil engineers must always take into consideration the influence of bond conditions on the overall cyclic behavior of beam–column joints to avoid a possible false interpretation of the pathology of earthquake-damaged RC structures. For instance, it is quite frequent after strong earthquakes to observe poorly detailed existing RC structures with almost (or even totally) undamaged beam–column joints, while the damage is limited solely to a particularly wide-open crack in the beams in the juncture with the joints. This type of damage may be mistakenly perceived by the inexperienced civil engineer as a result of ductile flexural failure and, hence, be insufficiently repaired during the retrofit process (i.e., be repaired solely by injecting epoxy resin). However, this failure does not by any means indicate the satisfactory hysteresis behavior of the joints. Contrarily, it conceals a particularly dangerous brittle failure mode: the pullout of the beam rebars from the joint region. This was clearly observed in a previous work of the authors [
11], where the seismic performance of exterior RC beam–column joint subassemblages with deficient straight anchorages of the longitudinal rebars at both the top and the bottom of the beam was investigated experimentally and analytically.
Another parameter which critically affects the performance of RC beam–column joints under cyclic lateral loading is the joint aspect ratio value, α =
hb/hc. In combination with the axial load of the column, the values of the joint aspect ratio, α, may cause the development of increased or decreased shear stresses in the beam–column joint region during an earthquake. In the experimental works of Meinheit and Jirsa [
12] and Parate and Kumar [
13], it was observed that the increase of the joint aspect ratio values resulted in an improved seismic performance, while Paulay and Park [
14] concluded that the beam–column joints of multi-story buildings with aspect ratio values greater than 1.0 are able to withstand significant shear stresses during cycling. Besides, the ratio α is related to the sectional dimensions of the connected structural members (beams and columns). Thus, an increase in the beam height,
hb, results in an increase in the ultimate shear capacity of the joint, as it is calculated according to the Tsonos model [
6,
15]. Of course, it should be noted that this increase in the beam height should be reasonable enough with respect to the column’s section height,
hc, and in accordance with the capacity design conception, which requires that the flexural strength ratio ΣM
Rc/ΣM
Rb should be greater than 1.30 [
3] or 1.20 [
16,
17]. Otherwise, the strong beam(s) may shift the seismic damage to the joint region or/and the columns. Meanwhile, the beam height,
hb, also affects the beam shear span/depth ratio value, a/d. For instance, a decrease in the beam height implies an increase in the beam shear span/depth ratio value, a/d, which, in turn, reduces the shear stresses developed in the beam. Thereby, a different seismic performance of the beam itself may be expected. The influence of the joint aspect ratio value, α, and of the value of the beam shear span/depth ratio, a/d, is even greater for poorly (non-seismically) detailed existing RC structures with respect to the modern ones. This is owed to the poor confinement provided to the joint region and the critical regions of beams and columns in combination with the low inherent strength and low quality of both concrete and steel.
The research found in the literature regarding the influence of the aforementioned parameters, especially when combined, in the hysteresis response of poorly detailed exterior RC beam–column joints is particularly limited. Kalogeropoulos and Tsonos experimentally investigated the seismic performance of original column subassemblages typical of substandard RC buildings built prior to the 1960–1970s with deficient lap splices, as well as the behavior of similar specimens retrofitted using thin steel jackets, RC jackets and CFRP jackets. They also proposed an analytical formulation for calculating the necessary confinement to achieve yielding of the deficiently lap-spliced rebars [
18,
19,
20]. Anagnostou et al. [
21] used a database with experimental results of 67 RC columns with lap-spliced steel bars, strengthened externally with FRP jackets subjected to pseudo-seismic loading to investigate the predictive performance of EC8.3 [
22] and KANEPE [
23] in calculating V
R and θ
u. In the experimental and analytical work of Sasmal et al. [
24], the seismic behavior of the exterior beam–column subassemblages of RC structure, designed and detailed according to the provisions of Eurocode and Indian Standards at different stages of their evolution (gravity load design, non-ductile, and ductile), was evaluated. Hakuto et al. [
25] performed seismic tests on interior and exterior RC beam–column joint subassemblages with a lack of or poor transverse reinforcement in the joint region, poor anchorage of the longitudinal bars passing through the interior joint, and beam rebar hooks bent out of the exterior joint core. Calvi et al. [
26] tested beam–column joint subassemblages typical of buildings designed in the 1950s and 1960s with plain steel rebars anchored with terminal hooks, without transverse reinforcement. They concluded that damage in the joints started at early stages, leading to hybrid failure mechanisms. In the experimental research of Zhao et al. [
27], the vulnerability of non-seismically detailed exterior connections with setback in columns was highlighted. The seismic behavior of non-seismically designed RC joints, regarding the effects of the beam–column depth ratio, column longitudinal reinforcement, and stirrups in joints on the seismic performance and shear strength of the joints, was investigated experimentally by Wong and Kuang [
28]. An empirical model on unreinforced beam–column RC joints with hook-ended plain bars was proposed by Teresa De Risi et al. [
29] to account for the peculiarities in terms of failure mode and concrete-to-steel interaction mechanisms. Beam tests and pull-out tests were performed by Fabrocino et al. [
30] to study in detail the force–slip relation of the bond mechanism for straight rebars and that of anchoring end details, i.e., circular hooks with a 180◦ opening angle. Verderame et al. [
31] experimentally investigated the seismic response of exterior unreinforced RC beam–column joints, representative of the existing non-conforming RC frame buildings, with different longitudinal reinforcements (plain or deformed), designed to be representative of two typical design practices (for gravity loads only or according to an obsolete seismic code). They observed different failure modes, namely joint failure with or without beam yielding. Recently, Karayannis et al. [
32] Naoum et al. [
33] and Karabini et al. [
34] tested full-scale beam–column joint subassemblages without stirrups in the joint region, while also using piezoelectric lead zirconate titanate (PZT) transducers for the examination of the efficiency of an innovative strengthening technique of RC columns and beam–column joints, which included external strengthening with carbon fiber-reinforced polymer (C-FRP) ropes. Fanaradelli et al. [
35] used pseudo-dynamic three-dimensional finite-element modeling to study the axial mechanical behavior of square and rectangular substandard RC columns, confined with fiber reinforced polymer (FRP) jackets and continuous composite ropes in seismic applications. The use of other innovative materials, such as glass fiber-reinforced polymer (GFRP) to strengthen existing structures, is also becoming more and more common [
36]. In the experimental work of Garcia et al. [
37], three full-scale substandard exterior RC beam–column joint specimens with inadequate detailing in the joint core zone were tested under cyclic loading and, subsequently, the damaged concrete of the joint was replaced by high strength concrete and the specimens were strengthened with CFRP sheets, showing a significant improvement in their hysteresis performance. Helal et al. [
38] investigated, experimentally and numerically, the seismic response of full-scale exterior RC beam–column joints with poor reinforcement details strengthened using post-tensioned metal straps for active confinement. A high ductile metal strap confinement, for the strengthening of low strength concrete columns, was used by Ijmai et al. [
39].
At this point, it is worth noting that the existing RC framed structures, built in the 1950–1970s period or earlier, possess numerous structural deficiencies; hence, their seismic response is dominated by the poor inelastic performance of the weaker structural members, namely the beam–column connections or/and the columns. The latter was clearly demonstrated many times in the aftermath of the strong earthquakes of the last 60 years, with the most recent examples being the seismic events in Turkey and Mexico in 2023. Furthermore, it cannot be ignored that the majority of RC structures worldwide were built prior to the imposition of modern code requirements for the design of earthquake-resistant RC structures. However, extending the service life of these structures and securing their ductile performance during future earthquakes through retrofitting processes are both rather challenging aims, but they also remain a more cost-effective and environmentally beneficial solution than demolishing and rebuilding. Thereupon, the satisfactory design of retrofit schemes, to allow for the strengthened structures to meet increased earthquake demands, requires a good understanding of the influences of various crucial design parameters in the seismic behavior of the beam–column joints, as well as of the developing failure mechanisms. Furthermore, the strengthening techniques and retrofit schemes should focus on combining measures exclusively undertaken for effectively improving bond conditions, for restoring the deficient anchorages, and for ensuring the load transfer and yielding of the lap-spliced column rebars, with the use of innovative materials and methods of application, which are less labor-demanding, easy to apply, cost-effective and environmentally friendly.
Along these lines, the present study aims to evaluate, both experimentally and analytically, the influence in the hysteresis behavior of critical parameters which are commonly combined in existing RC structures, namely the deficient anchorages of the beam longitudinal rebars in exterior beam–column joints, the inadequacy of the lap splices found in the columns, as well as the joint aspect ratio and beam shear span/depth ratio values. Intriguingly, the great number of design flaws and combined parameters examined herein allows for a better simulation of the actual behavior of substandard RC structures, while similar work found in the literature is extremely rare. Hence, the current experimental and analytical investigation provide a significant impetus to further the understanding of the overall hysteresis performance of existing structures, which can not be easily perceived when critical design parameters are examined separately.
2. Materials and Methods
Brittle failures are associated with catastrophic collapses of RC structures. Admittedly, there are multiple structural deficiencies, especially in the existing substandard structures built before the 1960–1970s, which trigger the development of brittle failure mechanisms in the most vulnerable members of the load bearing system, namely the columns and, particularly, the beam–column joints. Such shortcomings include the use of plain steel bars, the use of low compressive strength concrete, the inadequacy of anchorage and of the lap splice length of reinforcing bars, low confinement, etc. For this reason, an experimental program was conducted herein to evaluate the influence and profound implications of combined design defects on the seismic performance of substandard RC structures. After all, it is imperative to deeply understand how the hysteresis behavior is affected by these factors in order to effectively design the retrofit schemes for the earthquake-resistant rehabilitation of the existing RC structures.
Four exterior beam–column joint subassemblages with poor reinforcement details, representative of structural members found in existing substandard RC structures, were constructed and subjected to earthquake-type loading. The reinforcement details, cross-sectional dimensions and material properties of the specimens are summarized in
Table 1 and
Figure 1. Plain steel rebars with
were used as longitudinal reinforcement, while the transverse reinforcement consisted of plain steel ties with
. The concrete compression strength of the specimens was measured by using 150 × 300 mm cylinder compression tests (see
Table 1). The subassemblages were designated using a number, representing the number of the column longitudinal rebars, followed by the letters L (when lap splices of the column reinforcement exist), T (when the anchorage of the beam top rebars in the joint region is achieved with a 90° degree hook), B (when the anchorage of the beam bottom rebars in the joint region is achieved with a 90° degree hook) and one final letter (A or B) which corresponds to the shear span/depth of the beam ratio value (see
Table 1). In
Table 2, the nominal flexural moment capacity of the columns (
) and of the beam (
), the nominal shear capacity of the beam (
), and the capacity design ratio (
) for all specimens are summarized.
Figure 1.
Reinforcement details and cross-sections of beam–column joint subassemblages (a) 4TB-A, (b) 4LT-A, (c) 8T-A and (d) 8T-B.
Figure 1.
Reinforcement details and cross-sections of beam–column joint subassemblages (a) 4TB-A, (b) 4LT-A, (c) 8T-A and (d) 8T-B.
The seismic tests of the beam–column joint subassemblages 4TB-A, 4LT-A, 8T-A and 8T-B were conducted in the test setup shown in
Figure 2a,b, which is located at the Laboratory of Reinforced Concrete and Masonry Structures of the Aristotle University of Thessaloniki. The exterior RC beam–column joint subassemblages simulate the part of a real-world RC building between the points of contra-flexure, almost at the middle of the column height and the beam length, where the flexural moment from the seismic loading is near zero. This is depicted in
Figure 2b. The subassemblages did not include the RC slab. It should be noted that it is possible for beams at the perimeter of an RC building to experience torsional damage if significant stress is introduced into them by the slab. The inflection points of the columns were simulated using specific arrangements connected to the reaction frame and to the free ends of the columns by hinges. As a result, the vertical and the horizontal displacements of the columns’ ends were restrained, while still being able to rotate. The earthquake-type loading of the specimens was performed by subjecting the free end of the beam of each subassemblage to a large number of cycles of incremental inelastic lateral displacement amplitudes, according to the displacement-controlled schedule depicted in
Figure 3b, under a constant axial loading of the columns equal to 150 kN. A two-way actuator was used to apply the lateral displacements to the free end of the beam, while the resisted shear force was measured by a load cell (see
Figure 2a,b). The axial loading of the columns was achieved using a hydraulic jack. A calibrated linear variable differential transducer (LVDT) was used to measure the load point displacement. The steps of loading were determined using a test specimen similar to the examined ones, which was first loaded to its yield displacement. This was measured from the plot of resisted shear force-versus-displacement of the test specimen for the point when a significant decrease in stiffness occurred. This was further verified by the yielding of the longitudinal beam reinforcement at the column face. Thereupon, the loading was continued in the same direction (upper push half cycle) to 1.5 times the yield displacement, and the subassemblage was subsequently loaded in the opposite direction (upper and lower pull half cycle) to the same lateral displacement (see
Figure 3a). After the first cycle of loading, the maximum displacement of each subsequent cycle was incremented by 0.5 times the yield displacement [
25,
40,
41]. During the seismic tests, the subassemblages were subjected to a strain rate which corresponded to static conditions. As a result, the strengths exhibited by the specimens were somewhat lower than the strengths they would exhibit if subjected to load histories similar to actual seismic events, during which the strain rates are higher than the ones corresponding to static conditions [
42,
43,
44].
Table 1.
Reinforcement details and design parameters of the beam–column joint subassemblages.
Table 1.
Reinforcement details and design parameters of the beam–column joint subassemblages.
Specimen | Section Dimensions (mm) Column/ Beam | Longitudinal Reinforcement (mm) Column/ Beam | Transverse Reinforcement (mm) Column/ Beam | Anchorage of Beam Rebars | Lap Splices of Column Rebars (mm) | Shear Span/ Depth Ratio | fc (MPa) |
---|
4TB-A | 200 × 200/ 200 × 300 | 4Ø10/ 3Ø10 top 3Ø10 bottom | Ø6/200/ Ø6/200 | 90° hook (top) 90° hook (bottom) | - | 3.89 | 7.0 |
4LT-A | 200 × 200/ 200 × 300 | 4Ø10/ 3Ø10 top 3Ø10 bottom | Ø6/200/ Ø6/200 | 90° hook (top) Straight (bottom) | 200 | 3.89 | 7.0 |
8T-A | 200 × 200/ 200 × 300 | 8Ø10/ 3Ø10 top 3Ø10 bottom | Ø6/200/ Ø6/200 | 90° hook (top) Straight (bottom) | - | 3.89 | 8.0 |
8T-B | 200 × 200/ 200 × 200 | 8Ø10/ 3Ø10 top 3Ø10 bottom | Ø6/200/ Ø6/200 | 90° hook (top) Straight (bottom) | - | 6.18 | 8.0 |
Figure 2.
(a) Aerial view of the test setup and the instrumentation used; (b) reaction frame of the Laboratory of Reinforced Concrete and Masonry Structures of the Aristotle University of Thessaloniki where the seismic tests were performed, and the part of a RC structure which is simulated by the exterior beam–column joint subassemblages.
Figure 2.
(a) Aerial view of the test setup and the instrumentation used; (b) reaction frame of the Laboratory of Reinforced Concrete and Masonry Structures of the Aristotle University of Thessaloniki where the seismic tests were performed, and the part of a RC structure which is simulated by the exterior beam–column joint subassemblages.
Table 2.
Values of the nominal flexural and shear capacity, column shear force and capacity design ratio of the subassemblages.
Table 2.
Values of the nominal flexural and shear capacity, column shear force and capacity design ratio of the subassemblages.
| 4TB-A | 4LT-A | 8T-A | 8T-B |
---|
Nominal flexural capacity of the columns MRc (kNm) | (Over) 12.32 (Under) 11.20 | (Over) 12.32 (Under) 11.20 | (Over) 17.28 (Under) 16.64 | (Over) 17.92 (Under) 17.28 |
Nominal flexural capacity of the beam MRb (kNm) | 20.16 | 20.16 | 21.60 | 12.80 |
Nominal shear capacity of the beam VRb (kN) | 21.22 | 21.22 | 22.74 | 13.47 |
Capacity design ratio ΣMRc/ΣMRb | 1.17 | 1.17 | 1.57 | 2.75 |
Figure 3.
(a) Qualitative deformed shape of the specimens; (b) displacement-controlled schedule.
Figure 3.
(a) Qualitative deformed shape of the specimens; (b) displacement-controlled schedule.
4. Conclusions
An experimental and analytical investigation was conducted to evaluate the influence of critical design flaws, which are extremely common in existing substandard RC structures, on their seismic performance. Four subassemblages were designed with poor details to simulate the equivalent of structural members found in existing substandard RC structures. In particular, contrary to the provisions of modern codes for the design of earthquake-resistant RC structures (such as the Eurocode or ACI 318-19(22)), no consideration of the factors affecting the rebar anchorage was made (i.e., the location of the bar, the bar form, the provided concrete cover, the confinement provided by transverse reinforcement when welded to the rebar or not or by transverse pressure, or the bending diameter of bent reinforcing bars in the exterior RC beam–column joints). Furthermore, the ties used in the subassemblages were not closed with 135° hook-ends, as current building codes recommend. Instead, the ties of the subassemblages had 90° hook-ends, as used to be the practice during the 1950–1970s period. The lap splices of the longitudinal column reinforcing bars were of insufficient length in the case of subassemblage 4LT-A, as used to be the case in non-seismically designed RC structures. According to the provisions of the modern design codes, however, the design of col-umn lap splices of reinforcement must ensure load transfer between the spliced rebars under tension-compression reversals during seismic events, resulting in significantly increased lengths of lap splices.
Therefore, critical parameters, such as the use of plain steel bars and concrete with low compressive strength, the inadequacy (or adequacy) of anchorage of the beam bottom rebars in the joint region, the inadequacy of lap splice length of the column longitudinal reinforcement, and the joint aspect ratio and the shear span/depth of the beam ratio values, were examined. Based on the data acquired during testing, the interpretation of the results and the application of the proposed analytical formulation, the following conclusions are drawn.
The combination of design flaws and parameters examined herein critically affect the failure mode and the overall seismic response of the subassemblages, while allowing for the better simulation of the actual behavior of substandard RC structures. The latter can not be easily perceived when critical design parameters are examined separately. In particular:
The shear failure of the beam–column joint region, especially when unconfined, has a devastating impact on the seismic response of RC structures, being responsible for the disintegration of the joint concrete core and the consequential catastrophic collapse due to the loss of the axial load bearing capacity.
The inadequacy of the beam bottom rebar anchorage in the joint region results in premature bond-slip failure, excessive slipping and pullout of the bars. Hence, the shear forces inserted to the joint region from the beam bottom reinforcing bars is minimal, preventing shear failure from occurring in both diagonal directions of the joint. However, the hysteresis performance is particularly poor due to the excessive slipping of reinforcement (which is a brittle failure mode), and is dominated by the inability of the structure to dissipate seismic energy.
What is worth noting is, despite the concrete core of the joint in this case not being significantly disintegrated, the collapse of the structure can still occur due to the cumulative seismic energy under small deformations.
This also clearly demonstrates that the restoration of the beam rebar anchorage is imperative during the retrofit process to prevent the pullout of the beam rebars during a future earthquake and avoid the brittle response of the strengthened structure.
The inadequate lap splice length of the column rebars (especially when unconfined) further exacerbates the hysteresis performance, while preventing the development of the nominal flexural moment capacity of the column. Given that the lap splices are located in the potential plastic hinge region where a significant amount of seismic energy must be dissipated, the premature lap splice failure results in low plastic rotation capacity and possible collapse.
The column longitudinal reinforcement of subassemblages 8T-A and 8T-B was double the longitudinal reinforcement found in the columns of 4TB-A and 4LT-A. As a result, the nominal flexural moment capacity of the columns of 8T-A and 8T-B was increased with respect to that of subassemblages 4TB-A and 4LT-A. The same was true for the capacity design ratio values of subassemblages 8T-A and 8T-B, which were significantly higher than the corresponding values of 4TB-A and 4LT-A and with respect to the recommended value 1.30 (according to the Eurocode 8). However, this was not enough to ensure the ductile seismic response of the subassemblages, which exhibited poor hysteresis behavior dominated by the excessive slipping (pullout) of the beam bottom reinforcement. This further underlines the severe impact of the anchorage inadequacy in the seismic performance of existing poorly detailed RC structures.
Early bond-slip failure and premature slipping (or pullout) of the beam’s or/and column’s rebars prevent the development of the nominal flexural moment capacity of the beams or/and columns, while also significantly affecting the capacity design ratio value, the shear stresses inserted to the joint and, ultimately, the overall hysteresis behavior and failure mode of the beam–column joint. Eventually, the influence in the seismic behavior of both the lap splice length and the anchorage inadequacy can not be ignored when designing the retrofit schemes. Otherwise, the cyclic response of the strengthened structures may be overestimated and, hence, their structural integrity may be seriously jeopardized during future seismic events. Evidently, it is imperative to undertake specific measures exclusively for effectively improving bond conditions, for restoring the deficient anchorages, and for ensuring the load transfer and yielding of the lap-spliced column rebars. Moreover, based on the proposed analytical model, the design of the retrofit schemes should ensure that the beam–column joints of the strengthened structures will remain elastic during strong seismic excitations, while damage will be shifted and concentrated solely in the adjacent beams. This is possible when the calculated joint shear stress is less than half the joint shear capacity ().
For a RC beam–column connection with a rather low joint aspect ratio, α, despite demonstrating lower ultimate shear capacity with respect to a similar specimen with the same concrete compressive strength, , and a higher aspect ratio value, α, shear damage may not be developed due to the premature pullout of beam rebars, which dominates the failure mode, resulting in the formation of a significantly wide through crack in the beam. Thus, substantial local deformation of the well-anchored beam top rebars is owed to the dowel action, while minor shear forces are introduced to the joint region which remains intact.