Journal of Earthquake Engineering HPFRC Jacketing

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Journal of Earthquake Engineering Publication details, including instructions for authors and subscription information: http://www.tandfonline.com/loi/ueqe20

HPFRC Jacketing of Non Seismically Detailed RC Corner Joints a

a

b

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C. Beschi , P. Riva , G. Metelli & A. Meda a

Department of Engineering, University of Bergamo, Bergamo, Italy

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DICATAM, University of Brescia, Brescia, Italy

c

Civil Engineering Department, University of Rome “Tor Vergata,”, Rome, Italy Accepted author version posted online: 14 Aug 2014.Published online: 17 Nov 2015.

To cite this article: C. Beschi, P. Riva, G. Metelli & A. Meda (2015) HPFRC Jacketing of Non Seismically Detailed RC Corner Joints, Journal of Earthquake Engineering, 19:1, 25-47, DOI: 10.1080/13632469.2014.948646 To link to this article: http://dx.doi.org/10.1080/13632469.2014.948646

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Journal of Earthquake Engineering, 19:25–47, 2015 Copyright © A. S. Elnashai ISSN: 1363-2469 print / 1559-808X online DOI: 10.1080/13632469.2014.948646

HPFRC Jacketing of Non Seismically Detailed RC Corner Joints C. BESCHI1 , P. RIVA1 , G. METELLI2 , and A. MEDA3 Downloaded by [University of Brescia], [Giovanni Metelli] at 10:23 20 November 2014

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Department of Engineering, University of Bergamo, Bergamo, Italy DICATAM, University of Brescia, Brescia, Italy 3 Civil Engineering Department, University of Rome “Tor Vergata,” Rome, Italy 2

In this article, the effectiveness of a High-Performance Fiber-Reinforced Concrete (HPFRC) jacket for the seismic retrofitting of existing RC corner beam-column joints is experimentally investigated. The results of cyclic experimental tests on full-scale corner beam-column sub-assemblies (two unretrofitted and two retrofitted) are presented and discussed in detail, focusing on the effectiveness of the adopted retrofitting technique. The RC test units were designed with structural deficiencies typical of the Italian construction practice of the 1970s: use of smooth bars, inadequate reinforcement detailing, such as lack of stirrups in the joint panel, and hook-ended anchorage. The results underlined that the joint panel strength impairs the seismic response of the sub-structure with a drift at incipient collapse not greater than 2%. The results showed that the application of thin HPFRC jackets appears a promising technique to strengthen poorly-detailed RC joints: the jacket was able to increase the shear strength of the joint by about 40%, with respect to the bare joint, limiting the damage of the retrofitted sub-assembly, which reached an ultimate drift of 6%. Keywords Existing RC Structures; Seismic Retrofitting; HPFRC Jacketing; Corner Beam-Column Joints; Cyclic Test

1. Introduction The recent Italian earthquakes (Abruzzo 2009, Emilia Romagna 2012) dramatically demonstrated that a large amount of existing RC structures, designed for gravity loads only, were not able to sustain earthquake actions. This was mainly due to different structural deficiencies, which can be related to the absence of any capacity design principles, poor material properties, such as low-strength concrete, use of smooth bars for both longitudinal and transverse reinforcement, absence of transverse reinforcement in the joint regions, poor reinforcement detailing, such as insufficient amount of column longitudinal and transverse reinforcement, and inadequate anchorage detailing. Some of these construction details can be indicated as the potential critical causes of brittle failure mechanisms, which significantly reduce the overall ductility of the structure and lead to an inadequate lateral strength. For beam-column joints, three failure mechanisms can be identified. The first is related to the shear failure of the joint panel, typical of a weak-column/strong-beam system. This mechanism is brittle and, thus, it has to be avoided. The second one concerns the development of a plastic hinge in the beam, which is a desired ductile failure mode, typical of Received 21 January 2014; accepted 9 July 2014. Address correspondence to P. Riva, University of Bergamo v.le Marconi 5, 24044 Dalmine (BG), Bergamo, Italy. E-mail: [email protected] Color versions of one or more of the figures in the article can be found online at www.tandfonline.com/ueqe.

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a strong-column/weak-beam system. The third mechanism is due to the bond-slip failure mode, which depends on bond characteristics. These are related to the column size, which affects the bond length, to the type of reinforcement (smooth or deformed) and to the presence of transverse reinforcement in the joint, which may guarantee a confining action along the anchored bars. In this article, the attention will be focused on the seismic behavior of exterior beamcolumn joints. This kind of joint often represents the most critical regions in RC frames subjected to lateral loads, mainly owing to the absence of confinement in correspondence of at least one or two faces, the unbalanced thrust of the masonry infill, and a higher displacement demand caused by global torsional effects. In particular, the research has been focused on the behavior of corner beam-column joints with typical details of Italian construction practice in the 1960s and 1970s, which was characterized by the use of smooth bars with hooked-end anchorages. The brittleness of this type of external beam-column joint was firstly shown by tests carried out on a 2:3 scaled r.c. frame [Calvi et al., 2001]. The development of a shear failure mechanism markedly different from that provided in the case of a rigid joint behavior, for which a soft floor mechanism would be expected, was evident. Moreover, this failure mode was observed in several joints of poorly-detailed RC frames after recent Italian earthquakes [L’Aquila, 2009; Fig. 1]. In the literature, several experimental research works devoted to studying the seismic performance of r.c. frames designed for gravity loads only may be found. Many tests were carried out on sub-assemblies with interior or exterior beam-column joints characterized

FIGURE 1 Corner beam-column joint failure during the Abruzzo earthquake [ReLUIS, 2009]. © Dipartimento Protezione Civile (Reluis). Reproduced by permission of Dipartimento Protezione Civile (Reluis). Permission to reuse must be obtained from the rightsholder.


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by substandard reinforcing details. Most studies considered ribbed bars bent in the joint [Aycardi et al., 1994; Hakuto et al., 2000, Murty et al., 2003, Masi et al., 2013; Sharma et al., 2013; Sasmal et al., 2013] while few tests focused on sub-assemblies with hookedend smooth bars and only some of them were correctly designed to develop a joint shear failure [Calvi et al., 2001; Kam et al., 2010; Akguzel and Pampanin, 2008; Braga et al., 2009]. In fact, despite of the absence of transverse reinforcement in the panel region, some poorly detailed sub-assemblies [Russo and Pauletta, 2012; Masi et al., 2013] showed flexural hinges because a low reinforcement ratio was adopted in the beam, thus reducing the shear demand in the joint panel. In exterior beam-column joints with smooth bars and without transverse reinforcement in the joint panel, the shear transfer mechanism is based on a compression strut mechanism, whose efficiency depends on the concrete strength and on the anchorage solution adopted for the longitudinal beam reinforcement. If hooked anchorages were adopted, the joint strength would be impaired by the expulsion of a concrete wedge, due to the pushing action of the hooked-end anchorages in compression and caused by bar slip within the panel region [Calvi et al., 2001]. Because of the large amount in Italy of RC buildings designed for gravity loads only, before the introduction of adequate seismic design code provisions, the assessment of their seismic response has become an important and urgent issue, together with the development of repair and strengthening techniques. During the last decades, several techniques have been proposed for the seismic retrofit of RC elements [Fib Report, 1991; Fib Bullettin 24, 2003]. One of the most widespread techniques for upgrading RC elements is traditionally based on the use of steel plates, epoxy-bonded to the external surfaces of beams and slabs. As far as both cost and mechanical performance are concerned, this technique is simple and effective, but suffers from several disadvantages: corrosion of the steel plates, difficulty in handling heavy and long steel plates in tight construction sites, which are required in case of flexural strengthening of long girders. Concerning the strengthening of existing columns, the possibility of using RC jackets is usually considered, in particular when the elements are made of low strength concrete. Traditional jacketing presents some inconvenience, due to the jacket thickness being governed by the steel cover. This often leads to a jacket thickness higher than 70–100 mm, with a consequent significant increase of the column section geometry, which may alter the dynamic response of the structure [Bracci et al., 1995]. On the other hand, it is remarked that RC jacketing allows increasing not only the members’ strength but also its ductility, by offering an effective confining action [Campione et al., 2014]. However, concrete jacketing is probably the most labor-intensive strengthening method due to difficulties in placing additional joint transverse reinforcement. This method is successful in creating strong column-weak beam mechanisms, but suffers from considerable loss of floor space and disruption to the building occupancy. In Karayannis et al. [2008] a thin RC jacketing with small diameter reinforcements was proposed to repair damaged beam-column joints with bent-in deformed bars. The test results showed that the jacketing technique with dense reinforcements in the joint region allowed the sub-assembly to shift its failure mode form shear in the joint region to flexural hinging in the beam. Externally bonded FRP composites can eliminate some important limitations such as difficulties in construction and increase in member sizes [Gergely et al., 2000; Antonopoulos and Triantafillou, 2003; Mukherjee and Joshi, 2005; Alsayed et al., 2010]. The use of FRP offers several advantages, related to its high strength-to-weight ratio, resistance to corrosion, fast and relatively simple application. On the other hand, the risk of composite sheet delamination can impair the effectiveness of joint strengthening or repair,


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when anchoring mechanical devices are not adopted. Furthermore, FRPs may constitute a problem when fire resistance is a concern. Recently, a new technique based on the use of thin jackets made with HighPerformance Fiber-Reinforced Concrete (HPFRC) has been developed [Martinola et al., 2010]. This technique has been demonstrated to be effective for the strengthening of existing columns if compared with other techniques, particularly when the structure is made of low strength concrete [Beschi et al., 2011]. The proposed technique consists in encasing structural concrete elements in a thin layer (30–40 mm) of HPFRC material, which exhibits a weakly hardening behavior in tension coupled with a high compression strength, larger strain capacity and toughness when compared to traditional FRCs, which makes it ideal for use in members subjected to large inelastic deformation demands.

2. Experimental Investigation The present research work focused on experimental studies on the retrofitting of exterior beam-column sub-assemblies characterized by smooth bars with hooked-end anchorages. The sub-assemblies were strengthened by means of thin HPFRC jacketing with the aim to enhance the shear and bond-slip resistance of the joint by changing the undesirable brittle failure modes into a more ductile one, with the development of flexural hinges in the beams. In the following, the results of cyclic experimental tests on full-scale corner beamcolumn sub-assemblies (two unretrofitted and two retrofitted) are presented and discussed in detail, focusing on the effectiveness of the adopted retrofitting technique. 2.1. Geometry and Materials of Test Units 2.1.1 Unretrofitted Test Units. The unretrofitted test units, named in the following CJ1 and CJ2, were representative of a corner joint of the first level of a RC four-story frame designed according to the Italian design provisions in force before the 1970s [R.D., 1939] and suggested by the technical literature of that time [Santarella, 1945]. The reference building was faithful to the characteristics of the most part of the Italian building stock in the 1970s, such as frames in only one direction, usually the longitudinal one. The plan structural layout (10 x 21 m2 ) consisted of 5 bays in the longitudinal direction, with span of 4.5 m except for the central one which was 3 m long to place the stairs module, and two bays in the transverse direction with span of 5 m. The structural elements were designed only for gravity loads with the columns carrying only axial force and the beams designed according to the scheme of continuous beam on multiple supports with negative moments at the beam’s ends for compatibility reasons. As far as the materials proprieties are concerned, the steel grade of the reinforcement and the concrete strength are representative of materials adopted in the Italian buildings before the 1970s [Verderame et al., 2001; Verderame and Manfredi, 2001]. Further details of the design of the reference building may be found in Beschi [2012]. The beams were characterized by a 300 × 500 mm2 cross section, with smooth reinforcing bars with hooked-ends anchorages (Fig. 2). In the main beam 2 Ø12 and 2 Ø16 mm diameter longitudinal rebars were placed at the top and 2 Ø12 and 1 Ø16 mm diameter rebars were placed at the bottom with Ø8@200 mm stirrups, in order to avoid, with a relative high over-strength factor, any possible beam shear failure. In the transverse beam 2 Ø12 and 1 Ø16 mm upper bars and 2 Ø12 mm lower bars were adopted. The column cross section was 300 × 300 mm2 , with 4 Ø16 mm diameter longitudinal rebars. Lap splices with hook anchorages were adopted in the column longitudinal rebars. No transverse reinforcement was placed inside the joint, as was a common practice in the 1960s.



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FIGURE 2 Geometry and reinforcement details.

The size of the test units was determined by the distance between the contra-flexure points (assumed to be at mid-span of the beam and at mid-height of the column) for linear elastic lateral load response of the main longitudinal r/c frame in the reference building. The geometry of sub-assemblies and reinforcement details are shown in Fig. 2, where one may note the main longitudinal beam (2.1 m long) loaded by a shear force at the end and connected to the column (3.0 m high). It should be noticed that also a portion of the transverse secondary beam (0.65 m long) was realized, thus allowing a negative bending moment to be applied to the corner joint, in order to simulate the transverse action in the joint at service conditions (in the following all the details of the test set up are provided). The reinforcement had a mean yield strength (fym ) of 365 MPa and 445 MPa, respectively, for 12 and 16 mm diameter bars, evaluated by the results of three specimens tested according to EN 15630-3. The elongation at maximum tensile force (Agt ) was greater then 13.7% (Table 1). Normal grade concrete was supplied by a local ready-mix plant. The concrete was of medium workability (slump of about 150 mm according to EN 12350) with a maximum aggregate size of 16 mm. At time of testing, the average concrete compressive strength (fcm ) was 38.7 MPa, corresponding to C30/37 grade in accordance with ENV 1992-1-1:2004. This value was larger than what was considered in the design of the unretrofitted test units (C16/20), as shown in Riva et al. [2012] and Beschi [2012]. 2.1.2. Retrofitted Test Units. The retrofitting solution concerns the application of a HPFRC jacket to test units with the same geometry and detail of the unretrofitted ones (Fig. 2). After casting and a curing period of one month, the test units were prepared for the jacketing. To this end, the concrete surface was at first sandblasted, up to achieve a roughness of 1–2 mm, which had been demonstrated effective to ensure a good adhesion between new and old concrete even in the absence of chemical bonding agents [Martinola et al., 2010], and then watered to reach the saturation of the support (Fig. 3).

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TABLE 1 Material characteristics of steel reinforcement and HPFRC REINFORCEMENT φ


φ12 φ16 φ6 φ8

fym [MPa]

fum [MPa]

Agt [%]

365 445 493 337

558 546 556 440

15.9 13.7 16.1 21.0

HPFRC

Matrix

Steel fibers

fcm,cube [MPa]

ftm [MPa]

Ecm [GPa]

130

6.6

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fum [MPa]

leq [mm]

deq [mm]

Vf [%]

2000

15

0.18

1.2

φ: bar diameter; fym : average yield strength; fum : average ultimate strength; Agt : elongation at maximum tensile force - Tests carried out according to EN 15630-3 fum : mean ultimate tensile strength of wire; fcm,cube : average compressive strength; ftm : average direct tensile strength; Ecm : average elastic modulus; leq : equivalent length; deq : equivalent diameter; Vf : fibers volume

FIGURE 3 Specimens sandblasting (a), concrete surface before (b), and after sandblasting (c). The column was encased in a HPFRC jacket 40 mm thick while for the beam a U-shaped jacket was adopted with a thickness of 30 mm, which is the smallest value which may be adopted for technological limits. The jackets where cast in molds, with the specimen placed vertically, as in the reality. The strengthening material is a self-leveling mortar having a maximum aggregate size of 1.3 mm and water/binder (cement + microsilica) ratio equal to 0.17 by weight. The



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FIGURE 4 HPFRC characterization by means direct tensile test on dog-bone specimens.

mortar is reinforced with 1.2% (by volume) of straight steel micro-fibers having a length of 15 mm and a diameter of 0.18 mm. The ultimate tensile strength (fum ) of the wire is 2000 MPa. The compressive strength (fcm,cube ) of the HPFRC, as measured on 100 mm side cubes after 28 days of curing, was 130 MPa while the direct tensile strength (ftm ) was equal to 6.6 MPa (Table 1). Direct tensile test on dog-bone specimens were performed in order to characterize the material in tension. The experimental results, together with the specimen geometries, are reported in Fig. 4. As highlighted from the uniaxial tensile test, the material is characterized by a strain–hardening behavior in tension up to 0.15% strain, followed by a stable and slightly degrading softening behavior. The grade of the reinforcement was the same as for test units CJ1 and CJ2; the base concrete was characterized by an average compressive strength fcm of about 27 MPa. The retrofitted test units are labeled as RCJ1 and RCJ2 in the following.

3. Test Set-Up and Procedure The test set-up intended to reproduce the configuration of a corner beam-column subassembly of a frame subjected to reversed cyclic lateral loads acting on one of the two corner planes. To this aim, a test frame was designed to allow free-rotation both at the top and at the base of the column and to allow rotation and longitudinal translational motion of the main beam end to simulate the inflection points, which were assumed to occur at half of each element length (Fig. 5a). The test started with the application of an axial load equal to 210 kN to the column by means of two hydraulic jacks: the axial load was maintained constant during the entire test and represented the serviceability load acting on the column of the first level of the reference building evaluated according to the seismic load combination. A vertical force of 8.5 kN at the main beam’s end and a negative bending moment to the secondary beam were applied by means of hydraulic jacks, to simulate the combination of shear and moments acting in the joint in service conditions. The values of the moments applied to the joint were 18.7 kNm and 11.3 kNm at the main beam and secondary beam ends (including the effect of beam self- weight), respectively. The test was carried out with the application of cyclic loads in the main beam plane, by imposing at the top of the column cycles of displacements of increasing amplitude up to failure by means of an electromechanical jack, fixed to the reaction wall of the laboratory (Fig. 5b).


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FIGURE 5 Sub-assembly layout (a) and test set-up (b).

The loading history consisted of cycles characterized by drift increments referring to ACI protocol [ACI 437, 2007] with some adaptations to study the system at low drift levels, when the shear strength of the joint may impair the sub-assembly behaviour: 0.25% up to a drift of 1%, 0.5% up to a drift of 3% and 1% up to failure, as shown in Fig. 6a. Three fully reversed cycles were applied at each drift ratio. The test continued up to a drift ratio equal to 3% (90 mm top displacement) for the unretrofitted specimens and equal to 6% (180 mm top displacement) for the retrofitted specimens. In Fig. 6a, the experimental yield drifts for both directions of the unretrofitted sub-assembly are also shown. The yield drifts were evaluated by idealizing the experimental backbone curve as an elasto-plastic forcedeformation relationship with a secant stiffness at 75% of the ultimate load of the system [Park, 1988]. The yield drift allows for the appreciation of both the severity of the loading history and the number of plastic excursions during the cycling tests. In order to measure the horizontal displacements, potentiometric transducers were placed at the column top at the load application level (POS 1 and 2 in Fig. 6b). The rotations between the beams and the column were measured by means of a series of potentiometric transducers (POS 3-4-5-6 for the main beam and POS 7-8 for the secondary beam in Fig. 6b) while the rotations of the two halves of the column were monitored by the potentiometric transducers in POS 9-10-11-12 and POS 13-14-15-16, for the main and the secondary beam, respectively.

FIGURE 6 Loading history (a) and measurement devices (b).



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FIGURE 7 Horizontal load vs. horizontal displacement curves for test units CJ1 and CJ2. In order to gauge the opening of the cracks in the joint, potentiometric transducers were placed diagonally in the panel region for both the main beam (POS 17-18 in Fig. 6b) and the secondary beam (POS. 19-20). Furthermore, the transducers POS 21 and POS 22 measured the horizontal and the vertical displacements of the main beam end. Another transducer was provided to check any out of plane displacement. The horizontal load and the couple of forces applied to the secondary beam to reproduce the serviceability moment in the joint, were monitored by means of load cells, while the vertical load applied to the main beam was measured directly on the threaded bar placed at the beam’s end and was kept constant by means of a close loop hydraulic system.

4. Test Results 4.1. Unretrofitted Test Units The results in terms of horizontal load vs. displacement at the level of the load application point are shown for both the unretrofitted test units in Fig. 7. Due to the unsymmetrical amount of bar reinforcement at the top and bottom of the beam, the specimen clearly exhibited different lateral load responses in the positive and negative directions. It is observed that in the positive direction the beam lower face is in tension, whereas in the negative direction tension occurs in the beam upper face. Accordingly, as the cyclic loads were applied starting from the gravity service load condition, obtained by applying concentrated shear force acting downward at the main beam end, in the positive direction bending and shear actions in the main beam act in the opposite direction with respect to the initial ones, while in the negative direction they add up to the initially applied actions. In the positive direction, the specimens reached their maximum strength, equal to 31.3 kN for CJ1 test unit and 34.7 kN for CJ2 test unit at a drift equal to 2% and 2.5%, respectively. In the following loading cycles, the specimens exhibited a little strength degradation due to the formation of the expected flexural hinge in the main beam. At the final loading cycle, at 3% drift, the residual load carrying capacity was approximately 98% or


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96% of the maximum load for CJ1 and CJ2 test unit, respectively. In the second and third cycle at the same drift value, the specimens showed a reduction of the peak load equal to about 10% and 20%, respectively. In the negative load direction, the sub-assemblies response was governed by the shear damage in the joint panel. Both specimens achieved the maximum load at a 1% drift, equal to 35.98 kN and 35.41 kN for CJ1 and CJ2, respectively. After the peak value, the strength decreased more significantly for specimen CJ1 than for specimen CJ2 (63% and 76.5% of the peak value, respectively, at a drift equal to 3%). Also in the negative direction it was possible to recognize the trend observed for the positive one, with a lower maximum value of the load for the second and the third cycles of each sequence. At the limit state of incipient collapse, which may be observed at 1.5% and 2.0% for CJ1 and CJ2 specimen respectively, the strength reduction was about equal to 15% of the peak load (Fig. 7). The experimental results confirmed the high vulnerability of corner beam-column joints, characterized by a significant damage in the joint core. In addition, the pronounced cyclic stiffness degradation, with pinching effect in the hysteresis loops, showed the fundamental role played by bar-slip phenomena. In the positive direction the sub-assembly was characterized by the beam flexural failure, with a wide flexural crack at the interface with the joint, due to the slip of smooth reinforcing bars, which allowed a plastic behavior to be achieved. In the negative direction, the sub-assemblies were governed by the joint shear strength showing a rapid strength degradation beyond a 2% drift. The collapse occurred at 3% drift, and it was favored by the formation of an external concrete wedge due to end-hooks in compression, leading to a brittle local failure and a sudden loss of bearing capacity, as also observed by Calvi et al. [2001] on reduced-scale specimens. As shown in Fig. 8, which represents the evolution of the cracks pattern during the test, both test units showed early flexural cracks in the main beam, corresponding to a drift of 0.25% in the negative direction and 0.5% in the positive one, in agreement with the test set-up, which started with the application of a top-down vertical force at the beam’s end, causing a preliminary negative displacement of the sub-assembly. On the outer side of the joint, in correspondence with the secondary beam, no cracks appeared up to a drift of 0.75%, when the opening of a flexural crack at the bottom column-joint interface was observed. The first diagonal crack in the joint panel zone started in the negative direction during the first cycle at a drift equal to 1%. In the second positive cycle, at a drift equal to 2%, two diagonal cracks appeared in the opposite direction and the concrete wedge began to take shape. At a negative drift equal to 3% the joint strength reduction was greater than 40%. Severe cover spalling occurred in a wide area at the joint bottom (Figs, 8 and 9c). This wedge action was caused by the thrust of the hooked-end anchorages in compression and induced the plain bar slip within the panel joint. It is worth pointing out that no extensive flexural cracks and concrete crushing were observed in the inner side of the joint except at the main beam-joint vertical interface, due to the confining action of the secondary beam. In Fig. 9, the damage on the outer sides of test units CJ1 and CJ2 at the end of the test are shown. The three failure mechanisms are well highlighted: beam failure with a vertical crack at the beam-joint interface, joint shear failure with the diagonal cracks in the panel zone, and the thrust of the hooked-end beam bars in the column at the bottom of the joint. The diagram of the principal tensile stresses normalized with respect to the average cylindrical concrete strength versus drift is shown for test unit CJ1 in Fig. 10. The same trend was observed for test unit CJ2. It may be observed that at a drift equal to -1%, when

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FIGURE 8 Crack pattern in the outer sides of the joint for test units CJ1 and CJ2. the first √ diagonal crack appeared in the joint panel, the principal tensile stress is equal to 0.20 fcm for both un-retrofitted test units, confirming the results obtained by [Calvi et al., 2001], on the same kind of exterior beam-column joints. As expected, strength reduction occurred after cracking without any additional source for hardening behavior. The opening


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FIGURE 9 The damage at ±3% drift on the outer sides of the test units: CJ1 (a); CJ2 (b); and spalling due to the thrusting action of hooks in compression (c).

FIGURE 10 Principal tensile stress in the panel zone vs. drift curve for test units CJ1 and RCJ1. of a diagonal crack in the √ opposite direction was expected to happen for a principal tensile stress smaller than 0.2 fcm due to the cyclic damage. At a drift equal to +1%, a diagonal crack started to open in√the negative√direction, for which the maximum principal tensile stress was equal to 0.18 fcm and 0.2 fcm for CJ1 and CJ2 test unit respectively. It is worth pointing out that Italian Standard [D.M., 2008] and Eurocode 8 [CEN, 1998] √ suggested an unsafe value of the upper limit of the principal tensile stress (equal to 0.3 fcm ) for the assessment of a poorly-detailed joint strength. Finally, in Fig. 11, the deformation along the joint diagonal (in tension, mostly due to crack opening, while in compression due to concrete crushing) is shown, for increasing values of the drift applied to the specimens. At a drift equal to 2.0%, corresponding to a limit state of incipient collapse of the unreinforced sub-assemblies, a total shear crack width of about 2 mm was measured for both specimens. 4.2. Retrofitted Test Units Figures 12a and 12b plot the results in terms of horizontal load vs. displacement for the retrofitted test units. The shape of the curves is typical of the behavior of a section characterized by a RC core with a HPFRC jacket. The peak value corresponds to the achievement of the maximum tensile strength in the HPFRC jacket in the outer fibers of the main beam. In the positive direction, test unit RCJ1 reached its maximum capacity, equal to 44.5 kN in the first cycle at a drift equal to 0.75%, while for test unit RCJ2 the peak load was equal to 49.5 kN at the same drift. In both test units after the peak strength was reached,



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FIGURE 11 Deformations along the panel joint diagonals (devices 1 and 2) both for unretrofitted test units CJ1 and CJ2 and for retrofitted test units (RCJ1 and RCJ2).

FIGURE 12 Horizontal load Vc vs. drift curve for RCJ1 test unit (a) and for RCJ2 test unit (b). the load gradually decreased to the strength of the RC core-subassembly. The yielding of the bottom bar in the beam and the confinement ensured by the HPFRC jacket allowed to reach a high lateral drift. This behavior is evident from the plateaus in the curves for drift greater than +2%. In the negative direction the two test units reached approximately the same peak load of about 40 kN at a drift of –0.75%. After the peak, the shear load suddenly dropped and it leveled out around a value of 20 kN. This strength reduction, more evident in test unit RCJ1 than in the test unit RCJ2, was mainly due to the partial detachment of the HPFRC layer form the inner surface of the secondary beam core. As it can be noticed from Figs. 12a and 12b, both for positive and negative displacements, the strength reduction in the third cycle of each triplet at the same drift amplitude was about 15%−20%. With regard to the residual strength, test unit RCJ2 decayed slightly less than RJC1 in the positive direction (66% against 61% of test unit RCJ1) and slightly more for negative displacements (39% against 51%). In Fig. 13, the evolution of the cracks pattern for both reinforced specimens is depicted. It can be observed the formation at an early drift of a diagonal crack inside the joint panel, which didn’t develop significantly during the test, and a vertical flexural crack. The initial location of this vertical crack is different for the two specimens, being at the beam-joint interface for test unit RCJ1 and inside the joint for test unit RCJ2.

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FIGURE 13 Cracks pattern in the outer side of the joint for test units RCJ1 and RCJ2.



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The test units showed the first crack in the main beam in the negative direction in agreement with the test set-up, which started with the application of a top-down vertical force at the beam end, and as a consequence a preliminary negative moment acting on the beam. At the first cycle at –0.25% drift for both test units, also a diagonal crack appeared in the joint panel zone, which didn’t develop significantly in the following cycles. The crack width, starting from a value equal to 0.05 mm at a drift of –0.25%, increased up to a maximum value of 0.4 mm at a drift of –0.75%. In the following cycles, the damage also in the positive directions caused a partial closing of the diagonal crack up to a value of about 0.12 mm and 0.9 mm for the first and the second specimen respectively. The limited values of the crack opening in the joint panel are also plotted in Fig. 11 and compared to those of the unretrofitted specimens. For test unit RCJ1, at a drift equal to 0.75% a horizontal crack appeared in the column jacket at the bottom of the joint; also this crack didn’t develop as the test continued (Fig. 13). In the following cycles, for test unit RCJ1, the crack localized at beam-joint interface showed a significant opening increase up to values of about 45 mm at the end of the test (drift close to 6%). Only a few cracks of minor importance developed around this main crack. For test unit RCJ2, the damage localized at the beam-joint interface, with an increase of the crack width for positive moments while for negative moments, at a drift equal to -2%, a HPFRC wedge at the top of the joint began to spall off. Starting from a drift equal to 4% the crack spread upward along the jacket-column core interface. The damage pattern at the end of the test is also clearly shown in Fig. 14. No damage was observed on the outer side of the joint panel along the secondary beam direction. On the inner side of the secondary beam, the detachment of the HPFRC layer from the concrete support started to occur at a drift of about 1% for test unit RCJ1 (Fig. 14c). Test unit RCJ2 showed a HPFRC detachment limited to the jacket-column core interface for a drift greater than 3% (Fig. 14d).

FIGURE 14 Outer side of the main beam and of the joint in the retrofitted test units at the end of the test: (a) RCJ1; (b) RCJ2; detachment of the HPFRC jacket for test unit RCJ1(c); and RCJ2 (d).

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4.3. Comparison between Un-reinforced and Strengthened Test Units


It should be pointed out that the tested specimens were characterized by different concrete strength, equal to 38.7 MPa and 27 MPa for unretrofitted and retrofitted test units respectively. As a consequence, the comparison between the test results of the four test units may be performed only by means of dimensionless curves relating the applied drift to the column shear Vc normalized to the theoretical column shear Vc,th which causes the failure of the unretrofitted sub-assembly. The latter can be easily found by imposing the rotational equilibrium of the sub-structures (Fig. 5a) as: 1 ψMb,y , Vc,th = Lb Lbn Lc

(1)

where Lc = 3.0 m is the column height, Lb = 2.25 m is the distance between the column axis and the beam end, Lbn = 2.10 m the relative free span of the beam, Mb,y the beam moment resistance at the bar yielding, and Ψ is the ratio between the joint shear strength (Vjh ) and the shear value (Vjh,y ), related to the beam bending moment Mb,y by the following joint equilibrium equation: Vjh,y = Tb,y −

Lb Mb,y · , Lbn Lc

(2)

where Tby = Asi fymi is the tensile force in the reinforcement, and fymi is the bar yield strength. If Ψ is lower than 1, the joint shear strength Vjh governs the collapse of the subassembly (occurring for negative drift in the tests of unretrofitted specimens), while if Ψ is greater than 1 a plastic hinge can occur at the beam-joint interface (occurring for positive drift in the tests of unretrofitted specimens). The joint shear strength Vjh may be easily evaluated by means of the principal stress limitation model, proposed by Priestley [1997] and validated by several researchers [Pampanin et al., 2003; Celik and Ellingwood, 2008; Sharma et al. 2011, Riva et al., 2012]. The maximum tensile principal stress in the panel zone, corresponding to the development of the first diagonal crack during a first loading cycle, can be computed as follows: pt = k1 fcm ,

(3)

where fcm is the average cylindrical compressive strength of the concrete and k1 is a constant, calibrated on experimental results, with different values proposed in the literature and varying between 0.2 and 0.5 depending on the reinforcement detailing in the joint of exterior beam-column connections. In corner joints with hooked-end plain bars and without transverse reinforcements, the value proposed for k1 is equal to 0.2 [Calvi et al., 2001]. Assuming uniform normal and transverse stresses in the joint panel, the maximum resistant shear stress may be given by Mohr’s Circle, according to the following equation: vjh = pt 1 + fa pt ,

(4)

where fa = N/(bj hj ) = 2.33 MPa is the mean compressive stress on the column section due to the axial load N = 210 kN. The joint panel maximum shear strength can be calculated as follows:



41

Vjh = vjh hc hb ,

(5)

where hc = 300 mm and bb = 300 mm are the depth of the column and the width of the beam, respectively. Table 2 shows the main experimental results and the theoretical column shear action Vc,th causing the bare sub-assembly failure both for a concrete strength of 38.7 MPa (CJ1 and CJ2 test units) and for 27 MPa (RCJ1 and RCJ2 test units). Figure 15 plots the envelope curves of the normalized column shear (Vc /VC,th ) vs. the applied drift, thus allowing the evaluation of the effectiveness of the proposed technique by HPFRC jacketing. It can be noticed that the application of a HPFRC jacket may increase the peak shear strength of about 40% for positive displacements, and of about 50% for negative displacements. With respect to the residual strength, for both positive and negative directions the behavior of the retrofitted joints tended to the behavior of the unretrofitted ones, since for high drift the contribution of the HPFRC jacket was lost due to the detachment of the HPFRC jacket. Nevertheless, the strenghtend test units were able to withstand a drift up to 6%. This high lateral deformation was due to the benefical contribution of HPFRC jacketing which markedly reduced the shear damage in the joint panel and avoided the wedge concrete expulsion due to the thrust of the hooked-end anchorages in compression. The effectiveness of the adopted technique to improve the shear strength of exterior corner joints may be confirmed also by the value of √ the principal tensile √ stress in the joint panel. For RCJ1 test unit it reached a peak of 0.31 f√ cm and of 0.34 fcm for negative and positive drift, respectively, against the value of 0.20 fcm for un-retrofitted specimens (Fig. 10). Similar values have been found also for RCJ2 test unit. The normalized column shear stregth allows also to evaluate the effect of the transverse bending moment action in the joint due to the service load applied to the secondary beam. As shown in Table 2, the ratio between the maximum experimental shear action and the theoretical one is about 0.95 for both the unretrofitted sepcimens. This result highlights the low influence of action in the secondary beam on the joint strenght reduction (about 5%). A comparison between the experimental results of unretrofitted and retrofitted test units can be performed also in terms of dimensionless dissipated energy, calculated as the ratio between the dissipated energy Ei (hatched area in Fig. 16) and elastic energy E0i of the cycle with the same amplitude (see insert in Fig. 16). From the diagrams of Fig. 16, it can be noticed that the energy dissipation of specimens CJ1 and CJ2 are approximately comparable, with specimen CJ1 dissipating slightly more energy than specimen CJ2. The retrofitted specimens dissipated approximately 25% more energy than the unretrofitted ones at each drift value. However, unlike the unretrofitted sub-assemblies, for which the dissipated energy decreased starting from a drift equal to 2%, for the retrofitted RCJ1 test unit the energy dissipation always increased, because the HPFRC jacketing limited the joint damage even at high level of drifts (Fig. 17). For RCJ2 test unit the energy dissipation started to decline at 4% drift due to a more significant detachment of the HPFRC jacket with respect to RCJ1 test unit. Figure 17 shows the four test units at the end of the tests. For the unretrofitted test units CJ1 and CJ2 the three failure mechanisms could be clearly identified: beam failure with the vertical crack at the beam-joint interface, joint shear failure with the diagonal cracks in the panel zone, and the thrust of the hooked end beam bars in the column at the bottom of the joint (Figs. 17a and 17b). For the retrofitted test unit RCJ1, even if some thin cracks could be observed on the outer face of the joint next to the main beam, the damage localized mostly in the crack at the beam-joint interface (Fig. 17c). For test unit RCJ2 the vertical crack started a few centimeters inside the joint at the jacket-column core interface and developed upward only for high drift values (Fig. 17d).

42

CJ1 CJ2 RCJ1 RCJ2

Negative drift

2.00 2.50 0.75 0.75 −1.00 −1.50 −0.50 −0.50

−34.3 −34.7 −39.2 −40.2

dmax [%]

31.2 34.7 44.5 49.5

Vc,max [kN]

−36.4 −36.4 −27.4 −27.4

27.6 27.6 26.8 26.8

Vc,th

0.94 0.95 1.43 1.47

1.13 1.26 1.66 1.85

Vc,max Vc,th

−0.15 −0.57 −0.04 −0.04

0.47 0.58 0.03 0.05

−0.09 −0.34 −0.01 −0.00 0.44 0.98 0.28 0.32

w2 [mm]

w1 [mm]

−21.1 −26.3 −20.1 −15.6

30.4 33.2 26.8 32.8

Vc,r [kN]

−3.00 −3.00 −6.00 −6.00

3.00 3.00 6.00 6.00

du [%]

2.42 2.81 0.37 0.39

−1.51 −0.74 −0.02 0.08

wmax,1 [mm]

−1.51 −2.82 −0.04 −0.05

3.13 1.05 0.03 0.06

wmax,2 [mm]

Vc,max : maximum column shear; dmax : drift at Vc,max ; Vc,th : theoretical shear column causing the sub-assembly failure; w: deformation along the joint diagonal measured by devices 1 and 2 at peak Vc,max ; wmax : maximum shear crack width; du : ultimate drift

CJ1 CJ2 RCJ1 RCJ2

Positive drift

specimen

TABLE 2 Experimental results




43

FIGURE 15 Comparison between dimensionless envelope curves.

FIGURE 16 Dissipated energy.

For the retrofitted test units the diagonal crack in the joint panel appeared in the negative direction only at a drift equal to 0.25% with a width of about 0.06−0.07 mm, reaching a maximum value of 0.4 mm (Table 2). As previously described, the unretrofitted test units CJ1 and CJ2 showed significant diagonal crack opening, which reached values close to 3 mm.

5. Concluding Remarks The experimental results confirmed the seismic vulnerability of corner beam-column joints, designed with details typical of the Italian construction practice of the 1960s–1970s, characterized by the use of smooth bars with hooked-end anchorages and by the absence of


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FIGURE 17 The test units at the end of the tests: (a) CJ1; (b) CJ2; (c) RCJ1; and (d) RCJ2.

transverse reinforcement in the panel. The experimental results of two unretrofitted subassemblies showed a significant shear damage of the joint panel region and the slip of the beam bars in the joint, with the expulsion of a concrete wedge due to the thrust action of the hooks in compression. In addition, the joint strength seems not to be affected by the presence of a transverse action due to the service load acting on the secondary beam. Two sub-assemblies were strengthened by using a thin layer of HPFRC. The adopted technique is relatively simple since the material can be cast in a thin layer due to its selfleveling property. Experimental results demonstrated that this appears to be a promising technique. On the basis of the results discussed in this article, the following remarks can be drawn.

●

●

●

●

The application of a 30−40 mm thick HPFRC jacket on corner beam-column joint provides an increase of the normalized column shear of about 1.40 times with respect to the unretrofitted test units with a limited variation in the sub-assembly stiffness. The application of a HPFRC thin jacketing was able to shift the brittle joint shear failure to a more ductile beam flexural failure, according to the principles of capacity design. The damage of the retrofitted sub-assembly was limited to the joint-beam interface with diagonal crack width in the panel lower than 0.40 mm even at high drift level. The HPFRC jacket efficiently confined the joint avoiding the wedge concrete expulsion due to the thrust of the hooked-end anchorages. The proposed technique significantly improved the displacement capacity of the beam-column joint sub-assembly: the unretrofitted test units reached a drift equal to 3% against the 6% drift reached by retrofitted test units, value well beyond the specified code limits generally adopted for the ultimate limit state. Furthermore, the energy dissipation capacity of the retrofitted sub-assembly is up to 30% higher than the unretrofitted one, testifying a significant performance increase in case of seismic actions. Further research studies will be addressed to avoid the problem of the HPFRC detachment which led to a reduction of the post peak joint strength. The adoption of stud connectors between the host and the new concrete may be useful to control this phenomenon.


45

Funding The present work is part of the research supported by Re-LUIS, within the 2009–2012 project. The authors gratefully thank Tecnochem Italiana S.p.a., and Schnell S.p.a. for the financial and technical support to the research and Mr. Daniele Di Marco for the technical support in the experimental tests. The authors are grateful to engineer L. Bordoni for his assistance in carrying out the tests within his thesis work.


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Sharma, A., Eligehausen, R., and Reddy, G. R. [2011] “A new model to simulate joint shear behavior of poorly detailed beam–column connections in RC structures under seismic loads, part I: exterior joints,” Engineering Structures 33(3), 1034–1051. Sharma, A., Reddy, G. R., Vaze, K. K., and Eligehausen, R. [2013] “Pushover experiment and analysis of a full scale non-seismically detailed RC structure,” Engineering Structures 46, 218–33. Verderame, G. M., Stella, A., and Cosenza, E. [2001] “Mechanical properties of reinforcement used for r.c. constructions in ‘60s,” Proc. of the X National Conference ANIDIS - L’Ingegneria Sismica in Italia, September 9–13, Potenza-Matera (Italy), (in Italian). Verderame, G. M. and Manfredi, G. [2001] “Mechanical properties of concrete used for r.c. constructions in 60’s,” Proc. of the X National Conference ANIDIS - L’Ingegneria Sismica in Italia, September 9–13, Potenza-Matera (Italy), (in Italian).