Jacket Ing

8/18/2019 Jacket Ing

1/25

This article was downloaded by: [Bangalore University]On: 17 April 2015, At: 05:32Publisher: Taylor & FrancisInforma Ltd Registered in England and Wales Registered Number: 1072954 Registeredoffice: Mortimer House, 37-41 Mortimer Street, London W1T 3JH, UK

Click for updates

Journal of Earthquake EngineeringPublication details, including instructions for authors andsubscription information:

http://www.tandfonline.com/loi/ueqe20

HPFRC Jacketing of Non Seismically

Detailed RC Corner JointsC. Beschi

a, P. Riva

a, G. Metelli

b & A. Meda

c

a Department of Engineering, University of Bergamo, Bergamo, Italy

b 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 2014.

To cite this article: C. Beschi, P. Riva, G. Metelli & A. Meda (2015) HPFRC Jacketing of NonSeismically 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

PLEASE SCROLL DOWN FOR ARTICLE

Taylor & Francis makes every effort to ensure the accuracy of all the information (the “Content”) contained in the publications on our platform. However, Taylor & Francis,our agents, and our licensors make no representations or warranties whatsoever as tothe accuracy, completeness, or suitability for any purpose of the Content. Any opinionsand views expressed in this publication are the opinions and views of the authors,and are not the views of or endorsed by Taylor & Francis. The accuracy of the Contentshould not be relied upon and should be independently verified with primary sourcesof information. Taylor and Francis shall not be liable for any losses, actions, claims,proceedings, demands, costs, expenses, damages, and other liabilities whatsoever orhowsoever caused arising directly or indirectly in connection with, in relation to or arisingout of the use of the Content.

This article may be used for research, teaching, and private study purposes. Anysubstantial or systematic reproduction, redistribution, reselling, loan, sub-licensing,systematic supply, or distribution in any form to anyone is expressly forbidden. Terms &

http://crossmark.crossref.org/dialog/?doi=10.1080/13632469.2014.948646&domain=pdf&date_stamp=2014-08-14http://www.tandfonline.com/loi/ueqe20http://dx.doi.org/10.1080/13632469.2014.948646http://www.tandfonline.com/action/showCitFormats?doi=10.1080/13632469.2014.948646http://www.tandfonline.com/loi/ueqe20http://crossmark.crossref.org/dialog/?doi=10.1080/13632469.2014.948646&domain=pdf&date_stamp=2014-08-14


2/25

Conditions of access and use can be found at http://www.tandfonline.com/page/terms-and-conditions

http://www.tandfonline.com/page/terms-and-conditionshttp://www.tandfonline.com/page/terms-and-conditions


3/25

Journal of Earthquake Engi neering, 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

1Department of Engineering, University of Bergamo, Bergamo, Italy2DICATAM, University of Brescia, Brescia, Italy3Civil Engineering Department, University of Rome “Tor Vergata,” Rome, Italy

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 investi-gated. 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 effectivenessof the adopted retrofitting technique. The RC test units were designed with structural deficienciestypical of the Italian construction practice of the 1970s: use of smooth bars, inadequate reinforce-ment detailing, such as lack of stirrups in the joint panel, and hook-ended anchorage. The resultsunderlined 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 ableto 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 demon-

strated 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 deficien-

cies, 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 transversereinforcement, 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 devel-

opment 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.

25

mailto:[email protected]://www.tandfonline.com/ueqehttp://www.tandfonline.com/ueqehttp://www.tandfonline.com/ueqemailto:[email protected]


4/25

26 C. Beschi et al.

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 pres-

ence 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 beam-

column 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 dis-

placement 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 con-

struction 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, wasevident. 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.


5/25

HPFRC Jacketing of RC Corner Joints 27

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 hooked-

end 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 flex-

ural 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 gov-

erned 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 methoddue to difficulties in placing additional joint transverse reinforcement. This method is suc-

cessful 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, resis-

tance 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,


6/25

28 C. Beschi et al.

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 High-

Performance Fiber-Reinforced Concrete (HPFRC) has been developed [Martinola et al.,

2010]. This technique has been demonstrated to be effective for the strengthening of exist-

ing 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 beam-

column 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 andCJ2, 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 direc-

tion, 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 pro-

prieties 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 rein-forcing 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 rel-

ative 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.


7/25


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 onemay 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 ( f ym) of 365 MPa and 445 MPa, respectively,

for 12 and 16 mm diameter bars, evaluated by the results of three specimens tested accord-

ing 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 maximumaggregate size of 16 mm. At time of testing, the average concrete compressive strength ( f cm)

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 rough-

ness 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).


8/25

30 C. Beschi et al.

TABLE 1 Material characteristics of steel reinforcement and HPFRC

REINFORCEMENT

φ f ym [MPa] f um [MPa] Agt [%]

φ12 365 558 15.9

φ16 445 546 13.7

φ6 493 556 16.1

φ8 337 440 21.0

HPFRC

f cm,cube [MPa] f tm [MPa] E cm [GPa]

Matrix 130 6.6 36

f um [MPa] leq [mm] d eq [mm] V f [%]

Steel fibers 2000 15 0.18 1.2

φ: bar diameter; f ym: average yield strength; f um: average ultimate strength; Agt : elongation atmaximum tensile force - Tests carried out according to EN 15630-3 f um: mean ultimate tensile strength of wire; f cm,cube: average compressive strength; f tm: average directtensile strength; E cm: average elastic modulus; leq: equivalent length; d eq: equivalent diameter; V f :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


9/25


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 ( f um) of the wire is

2000 MPa. The compressive strength ( f cm,cube) of the HPFRC, as measured on 100 mm side

cubes after 28 days of curing, was 130 MPa while the direct tensile strength ( f tm) 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 f cm of about 27 MPa. Theretrofitted 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 subassem-

bly 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).


10/25

32 C. Beschi et al.

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 force-

deformation 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 potentiomet-

ric 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).


11/25


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 repro-

duce the serviceability moment in the joint, were monitored by means of load cells, whilethe 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 thebeam, 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 ten-

sion, 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 posi-

tive 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 degra-

dation 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


12/25

34 C. Beschi et al.

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 fail-

ure, 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 direc-

tion, 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’send, 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


13/25


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√ f cm 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


14/25

36 C. Beschi et al.

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√

f cm due to the cyclic damage. At a drift equal to +1%, a diagonalcrack started to open in the negative direction, for which the maximum principal tensile

stress was equal to 0.18√

f cm and 0.2√

f cm 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√

f cm) 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 increasingvalues 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 charac-

terized 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,


15/25


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 V c 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 andit 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 theresidual 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.


16/25

38 C. Beschi et al.

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


17/25


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 col-

umn 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 interfaceshowed 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 theend of the test: (a) RCJ1; (b) RCJ2; detachment of the HPFRC jacket for test unit RCJ1(c);

and RCJ2 (d).


18/25

40 C. Beschi et al.

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 respec-

tively. 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 col-umn shear V c normalized to the theoretical column shear V c,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:

V c,th = Lb

Lbn 1

Lcψ M b, y, (1)

where Lc = 3.0 m is the column height, Lb = 2.25 m is the distance between the columnaxis and the beam end, Lbn = 2.10 m the relative free span of the beam, M b,y the beammoment resistance at the bar yielding, and Ψ is the ratio between the joint shear strength

(V jh) and the shear value (V jh,y), related to the beam bending moment M b,y by the following

joint equilibrium equation:

V jh, y = T b, y − Lb

Lbn· M b, y

Lc, (2)

where T by = Asi f ymi is the tensile force in the reinforcement, and f ymi is the bar yieldstrength.

If Ψ is lower than 1, the joint shear strength V jh governs the collapse of the sub-

assembly (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 posi-tive drift in the tests of unretrofitted specimens). The joint shear strength V jh 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 devel-

opment of the first diagonal crack during a first loading cycle, can be computed as

follows:

pt = k 1

f cm, (3)

where f cm is the average cylindrical compressive strength of the concrete and k 1 is a con-

stant, 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 k 1 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:

v jh = pt

1 + f a

pt , (4)

where f a = N /(b jh j) = 2.33 MPa is the mean compressive stress on the column section dueto the axial load N = 210 kN.

The joint panel maximum shear strength can be calculated as follows:


19/25


V jh = v jhhchb, (5)

where hc = 300 mm and bb = 300 mm are the depth of the column and the width of thebeam, respectively.

Table 2 shows the main experimental results and the theoretical column shear action

V c,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 (V c/V C,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 displace-

ments. 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 driftup 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√

f cm for negative and posi-

tive drift, respectively, against the value of 0.20√

f cm 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 E i (hatched area in Fig. 16) and elastic energy E 0i 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 dissi-pated 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).


20/25

T A B L E

2

E x p e r i m e n t a l r e s u

l t s

s p e c

i m e n

V c , m a x

[ k N ]

d m a x

[ % ]

V c , t

h

V c , m

a x

V c , t

h

w 1

[ m m ]

w 2

[ m m ]

V c , r

[ k N ]

d u

[ % ]

w m a x , 1

[ m m

]

w m a x , 2

[ m m ]

P o s i t i v e

d r i

f t

C J 1

3 1 . 2

2 . 0

0

2 7 . 6

1 . 1 3

− 0

. 0 9

0 . 4

7

3 0

. 4

3 . 0

0

− 1

. 5 1

3 . 1

3

C J 2

3 4 . 7

2 . 5

0

2 7 . 6

1 . 2 6

− 0

. 3 4

0 . 5

8

3 3

. 2

3 . 0

0

− 0

. 7 4

1 . 0

5

R C J 1

4 4 . 5

0 . 7

5

2 6

. 8

1 . 6 6

− 0

. 0 1

0 . 0

3

2 6

. 8

6 . 0

0

− 0

. 0 2

0 . 0

3

R C J 2

4 9 . 5

0 . 7

5

2 6

. 8

1 . 8 5

− 0

. 0 0

0 . 0

5

3 2

. 8

6 . 0

0

0 . 0

8

0 . 0

6

N e g a t i v

e d r i

f t

C J 1

− 3 4 . 3

− 1

. 0 0

− 3 6

. 4

0 . 9 4

0 . 4

4

− 0

. 1 5

−

2 1

. 1

− 3

. 0 0

2 . 4

2

− 1

. 5 1

C J 2

− 3 4 . 7

− 1

. 5 0

− 3 6

. 4

0 . 9 5

0 . 9

8

− 0

. 5 7

−

2 6

. 3

− 3

. 0 0

2 . 8

1

− 2

. 8 2

R C J 1

− 3 9 . 2

− 0

. 5 0

− 2 7 . 4

1 . 4 3

0 . 2

8

− 0

. 0 4

−

2 0

. 1

− 6

. 0 0

0 . 3

7

− 0

. 0 4

R C J 2

− 4 0 . 2

− 0

. 5 0

− 2 7 . 4

1 . 4 7

0 . 3

2

− 0

. 0 4

−

1 5

. 6

− 6

. 0 0

0 . 3

9

− 0

. 0 5

V c , m a x : m

a x

i m u m c o

l u m n s h e a r ; d m a x :

d r i

f t a t V c , m a x ; V c , t

h : t h e o r e t i c a

l s h e a r c o l u

m n c a u s i n g t h e s u

b - a s s e m

b l y f a i l u

r e ; w :

d e

f o r m a t i o n a

l o n g t h e

j o i n t d

i a g o n a

l m e a s u r e

d

b y

d e v

i c

e s

1 a n

d 2 a t p e a

k V c , m a x ; w m a x : m a x

i m u m s h e a r c r a c

k w

i d t h ; d u : u

l t i m a t e d r i

f t

42


21/25


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 nega-

tive direction only at a drift equal to 0.25% with a width of about 0.06−0.07 mm, reachinga 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, char-

acterized by the use of smooth bars with hooked-end anchorages and by the absence of


22/25

44 C. Beschi et al.

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 sub-

assemblies 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 self-

leveling 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 jointprovides 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 capac-

ity 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 highdrift 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.


23/25


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.

References

American Concrete Institute [2007] ACI 437.1R-07: Load Tests of Concrete Structures: Methods,

Magnitude, Protocols, and Acceptance Criteria, American Concrete Institute, Farmington Hills,

Michigan.

Akguzel, U. and Pampanin, S. [2008] “Effects of variation of axial load and bi-directional loading on

the FRP retrofit of existing B-C-joints,” The 14th World Conference on Earthquake Engineering,

Beijing, China.

Alsayed, S. H., Al-Salloum, Y. A., Almusallam, T. H., and Siddiqui, N. A. [2010] “Seismic responseof FRP-upgraded exterior RC beam-column joints,” Journal of Composites for Construction

14(2),195–208.

Antonopoulos, C. P. and Triantafillou, T. C. [2003] “Experimental investigation of FRP-strengthened

RC beam-column joints,” Journal of Composites for Construction 7(1), 39–49.

Aycardi, L. E., Mander, J. B., and Reinhorn, A. M. [1994] “Seismic resistance of reinforced concrete

frame structures designed only for gravity loads: experimental performance of subassemblages,”

ACI- Structural Journal 91(5), 552–563.

Beschi, C. [2012] “Retrofitting of beam-column joints under sesimic loads,” Ph.D. Thesis, University

of Brescia, Brescia, Italy.

Beschi, C., Meda, A., and Riva, P. [2011] “Column and joint retrofitting with high performance fiber

reinforced concrete jacketing,” Journal of Earthquake Engineering 15(7), 989–1014.Bracci, J. M., Reinhorn, A. M., and Mander, J. B. [2010] “Seismic resistance of reinforced concrete

frame structures designed for gravity loads: performance of structural system,” ACI Structural

Journal 92(5), 597–608.

Braga, F., Gigliotti, R., and Laterza, M. [2009] “Existing structures with smooth reinforcing

bars: experimental behaviour of beam-column joints subject to cyclic lateral loads,” The Open

Construction & Building Technology Journal 3, 52–67.

Calvi, G. M., Magenes, G., and Pampanin, S. [2001] “Studio sperimentale sulla risposta sismica di

edifici a telaio in cemento armato per soli carichi da gravità,” Proc. of the X National Conference

ANIDIS - L’Ingegneria Sismica in Italia, September 9–13, Potenza-Matera, Italy.

Campione, G., Fossetti, M., Giacchino, C., and Minafò, G. [2014]. “RC columns externally

strengthened with RC jackets,” Materials and Structures, 47(10), 1715–1728.

Celik, O. C. and Ellingwood, B. R. [2008] “Modeling beam-column joints in fragility assessmentof gravity load designed reinforced concrete frames,” Journal of Earthquake Engineering 12,

357–381.

Circolare 2 Febbraio 2009 n.617 [2009] “Instructions for the application of the new technical code

for structures (In Italian),” G.U. n.47, February 26, 2009, Supplemento Ordinario n.27, Rome.

Dipartimento Protezione Civile, ReLUIS [2009] “Draft guidelines for the restoration and strength-

ening of structural elements, infill and partition walls” (in Italian), Dolce, M. and Manfredi, G.

(eds.). ISBN: 978-88-89972-29-8.

D.M. 14 Gennaio 2008 [2008] “Technical Code for Structures (in Italian),” G.U. n.29, February 2

2008, Supplemento Ordinario n.30, Rome.

EN 1992-1-1 [2004] Eurocode 2: Design of Concrete Structures - Part 1-1: General Rules, and Rules

for Buildings, European Committee for Standardization, Brussels.


24/25

46 C. Beschi et al.

EN 1998-1-1 [2004] Eurocode 8: Earthquake Resistant Design of Structures—Part 1: General Rules

and Rules for Buildings, CEN, Brussels.

EN 12350-2 [2001] Testing Fresh Concrete – Slump Test, European Committee for Standardization,

Brussels.

EN 15630-1 [2010] Steel for the Reinforcement and Prestressing of Concrete. Test Methods - Part 1:

Reinforcing Bars, Wire Rod and Wire, European Committee for Standardization, Brussels.Fib Bulletin No. 24. [2003] “Seismic assessment and retrofit of reinforced concrete buildings,” State-

of-art report prepared by Task Group 7.1, The International Federation for Structural Concrete,

August 2003, 312 pp.

Fib Report [1991] “Repair and strengthening of concrete structures. Guide to good practice.” Thomas

Telford Ltd. 1991, 37 pp, London.

Gergely, J., Pantelides, C. P., and Reaveley, L.D. [2000] ”Shear strengthening of RCT-joints using

CFRP composites,” Journal of Composites for Construction 4(2), 56–64.

Hakuto, S., Park, R., and Tanaka, H. [2009] “Seismic load tests on interior and exterior beam-column

joints with substandard reinforcing details,” ACI Structural Journal 97(1), 11–25.

Kam, W. Y., Quintana Gallo, P., Akguzel, U., and Pampanin, S. [2010] Influence of slab on the

seismic response of sub-standard detailed exterior reinforced concrete beam column joints. 9th USNational and 10th Canadian Conference on Earthquake Engineering: Reaching Beyond Borders,

Toronto, Canada, July 25–29, 2010.

Karayannis, C. G., Chalioris, C. E., and Sirkelis, G. M. [2008] “Local retrofit of exterior RC beam-

column joints using thin RC jackets - An experimental study,” Earthquake Engineering and

Structural Dynamics 37(5), 727–746.

Martinola, G., Meda, A., Plizzari, G. A., and Rinaldi, Z. [2010] “Strengthening and repair of RC

beams with fiber reinforced concrete,” Cement and Concrete Composites 32(9), 731–739.

Masi, A., Santarsiero, G., and Nigro, D. [2013] “Cyclic tests on external RC beam-column joints:

role of seismic design level and axial load value on the ultimate capacity,” Journal of Earthquake

Engineering 17(1), 110–136.

Mukherjee, A. and Joshi, M. [2005] “FRPC reinforced concrete beam-column joints under cyclic

excitation,” Composite Structures 70(2), 185–199.Murty, C. V. R., Rai, D., Bajpai, K. K., and Jain, S. K. [1994] “Effectiveness of reinforcement details

in exterior reinforced concrete beam–column joints for earthquake resistance,” ACI Structural

Journal 100(2), 149–56.

Pampanin, S., Magenes, G., and Carr, A. [2003] “Modeling of shear hinge mechanism in poorly

detailed RC beam-column joints,” Proc. of the FIB 2003 Symposium, May 6–8, Athens, Greece.

Park, R. [1988] “State-of-the art report: ductility evaluation from laboratory and analytical testing,”

Proc. 9th WCEE. IAEE , Tokyo–Kyoto, Japan VIII, pp. 605–616.

Priestley, M. J. N. [1997] “Displacement-based seismic assessment of reinforced concrete buildings,”

Journal of Earthquake Engineering 1(1), 157–192.

Regio Decreto R.D. 16/11/1939 n. 2229 [1940] “Norme per la esecuzione delle opere in conglom-

erato cementizio semplice e armato,” G.U. n. 92, 18/04/1940.Riva, P., Beschi, C., Metelli, G., and Messali, F. [2012] “Modelling of exterior beam to column joint

in ’60-70s RC existing buildings,” Proc. of 4th International Symposium: Bond in Concrete 2012:

Bond, Anchorages, Detailing 1, 225–232.

Russo, G. and Pauletta, M. [2012] “Seismic behavior of exterior beam-column connections with plain

bars and effects of upgrade,” ACI Structural Journal 109(2), 225–234.

Santarella, L. [1945] “Il cemento armato – Le applicazioni alle costruzioni civili e industriali,” Milano

Edizioni Hoepli.

Sasmal, S., Ramanjaneyulu, K., Balthasar, N., and Lakshmanan, N. [2013] “Analytical and exper-

imental investigations on seismic performance of exterior beam–column subassemblages of

existing RC-framed building,” Earthquake Engineering & Structural Dynamics 42, 1785–1805.


25/25


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 usedfor 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. con-

structions in 60’s,” Proc. of the X National Conference ANIDIS - L’Ingegneria Sismica in Italia,

September 9–13, Potenza-Matera (Italy), (in Italian).

Jacket Ing

Documents

Transcript of Jacket Ing