Results from cyclic tests on high aspect ratio RC columns strengthened with FRP systems

Construction and Building Materials 37 (2012) 606–620

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Results from cyclic tests on high aspect ratio RC columns strengthened with FRP systems Roberto Realfonzo ⇑, Annalisa Napoli Department of Civil Engineering University of Salerno, Via Ponte don Melillo, 84084 Fisciano (SA), Italy

h i g h l i g h t s " Experimental tests on full scale RC 300  700 mm rectangular columns are presented. " Performances of different strengthening techniques with FRP materials are discussed. " It is shown that the FRP confinement significantly increases the columns ductility. " The combined use of FRP confinement and steel angles also improves the flexural strength. " The FRP confinement increases the cumulative dissipated energy of columns.

a r t i c l e

i n f o

Article history: Received 28 February 2012 Received in revised form 5 July 2012 Accepted 22 July 2012 Available online 7 September 2012 Keywords: Confinement RC columns FRP Ductility Experimental results

a b s t r a c t A limited number of experimental studies have been performed on FRP-confined columns having rectangular cross section with high aspect ratio. The experimental study presented herein addresses this knowledge gap by investigating performances, under constant axial load and cyclically reversed horizontal force, of full scale rectangular (300  700 mm) RC columns externally confined by using Fiber Reinforced Polymers (FRPs) or strengthened with FRP wraps and steel profiles. The study is a part of a wider experimental campaign also including tests on square (300  300 mm) columns whose details have been already published elsewhere. Test results discussed herein aim to evaluate the influence on the column performance of relevant parameters, such as: unconfined concrete strength, longitudinal steel reinforcement (smooth or deformed rebars), strengthening system (FRP confinement with or without steel profiles); axial load level and number of FRP layers. Finally, the effectiveness of the used strengthening techniques is investigated through the comparison with the performances obtained in the case of square members. Ó 2012 Elsevier Ltd. All rights reserved.

1. Introduction Over the past two decades, the demand for seismic repair and retrofitting of existing reinforced concrete (RC) structures is significantly increased. It is known that the most building heritage built prior to the 1970s was designed in order to withstand only gravity loads or according to outdated seismic rules; thereby, undesired collapse mechanisms due to irregular distribution of strength and stiffness and/or to a low rotational ductility of some structural components have been frequently observed under seismic events. In particular, these ‘‘under designed’’ structures are often characterized by an unsatisfactory structural behavior due to a weak column-strong beam assembly that, under a seismic event, yields

⇑ Corresponding author. Tel.: +39 089 964085; fax: +39 089 968739. E-mail address: [email protected] (R. Realfonzo). 0950-0618/$ - see front matter Ó 2012 Elsevier Ltd. All rights reserved. http://dx.doi.org/10.1016/j.conbuildmat.2012.07.065

most likely to the formation of local hinges in the columns and to a consequent low available global ductility. In fact, the column hinging is frequently associated to the occurrence of a local collapse mechanism with high ductility demand for few structural components involved in such mechanism. Conversely, the available ductility of columns is limited (significantly lower than that in beams) as their deformation capacity is penalized by the presence of the axial force, crushing of the concrete in compression and instability of the steel reinforcing bars. In existing buildings, these last two phenomena often occur due to the inadequate size, spacing and end-closure of stirrups. In order to improve the strength and manly the ductility of under-designed RC columns, external confinement systems employing Fiber Reinforced Polymer (FRP) materials have emerged as a promising alternative to the traditional strengthening techniques, such as steel or concrete jacketing. The use of FRP confining systems does not allow, except for particular cases, to convert the local collapse mechanism in a global

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one (strong column-week beam behavior); however, it assures a greater availability of global ductility by increasing the local one. The growing interest in FRP material is confirmed not only by the increasing number of practical applications, but also by the development of some national and international codes regarding this area, such as: the ACI440 [1] in USA; the fib bulletin No. 14 [2] in Europe; and the DT200 [3] in Italy edited by the National Research Council (C.N.R.). So far, the literature on FRP confined RC elements is vast but it has mainly regarded the study of results from compression tests performed on small scale specimens having different cross sections and confined with systems (wet lay-up systems or FRP tubes) made of carbon (CFRP), glass (GFRP) or aramid (AFRP) fibers. More recent is the use of natural (basalt, hemp, etc.) and steel (SRP) fibers. Additionally, several analytical models able to predict the compressive strength, the corresponding ultimate axial strain and the stress–strain constitutive law of the FRP confined concrete have been developed; advanced states of art on the mentioned topics can be found in [4,5]. However, if the behavior of small concrete specimens under compression has been extensively investigated, a relatively limited number of experimental studies have been performed on full scale FRP confined RC columns subjected to axial load and cyclic flexure [6–8], and only some studies accounted for the presence of smooth steel rebars as longitudinal reinforcement of the tested columns [9,10]. Furthermore, even more limited are the results available from tests performed on FRP-confined columns having rectangular cross section with high aspect ratio [8,11,12]. It is noted that the Italian Guidelines DT200 [3] discourage the use of FRP confinement for members with rectangular cross-section having a side ratio greater than 2 if not otherwise proven by suitable experimental tests; this because such applications are not carefully validated by testing. The experimental study presented herein addresses these knowledge gaps by investigating the performance of nine full scale RC columns, with 300  700 mm rectangular cross section (i.e. aspect ratio of about 2.3), having longitudinal reinforcement made of smooth or deformed steel rebars. These columns, representative of existing structural members with inadequate seismic details, were tested under constant axial load and cyclically reversed horizontal force at the Testing Laboratory for Materials and Structures of the University of Salerno, Italy. Of these members: five were externally confined with two or four CFRP layers; two were strengthened using, in addition to the CFRP jacket, four longitudinal steel profiles placed in the corners of the columns; the remaining were not reinforced and used as ‘‘control samples’’. The preliminary results of these tests can be found in Napoli et al. [13]. It is worth noting that the nine tested members are included in a larger experimental program consisting of 33 tests, of which 24 were performed on columns having a 300  300 mm cross section. Results referring to the 24 square members, omitted herein for the sake of brevity, are extensively discussed in Realfonzo and Napoli [14]. Test results have allowed evaluating the benefits introduced by the considered strengthening systems in terms of strength and ductility by varying some relevant parameters, such as: the typology of the longitudinal steel reinforcement (smooth or deformed rebars); the concrete strength; the type of retrofitting system (FRP confinement; FRP confinement plus steel profiles); the number of CFRP layers used for column wrapping (two or four layers); and the level of applied axial load (two different levels). The effect of the FRP confinement on the crack pattern and failure mode of test specimens was also investigated.

607

Finally, comparisons with some relevant results obtained from tests on square columns were examined, and the efficiency of the FRP confinement depending of the shape of the cross-section (square and rectangular) was discussed. 2. Experimental campaign 2.1. Test specimens and strengthening systems A total of nine full scale RC columns were tested at the Testing Laboratory for Materials and Structures of the University of Salerno under a constant axial load and a cyclically reversed horizontal force. The specimens have a rectangular 300  700 mm2 cross section, a length of 2500 mm and a concrete foundation of dimensions 1400  600  800 mm. Details about the specimen configurations (geometry, reinforcements, etc.) are illustrated in Fig. 1. As mentioned earlier, these specimens are included in a larger test matrix comprising twenty-four 300  300 mm members, 2200 mm long, whose results have been already published in [14]. All specimens were designed and realized to be representative of structural components belonging to gravity load designed (GLD) existing buildings. For this reason, a low strength concrete was selected: the concrete mixture was studied in order to obtain a mean value of the cylindrical compression strength (fcm) of about 14 MPa. The concrete strength per each column was estimated by testing in compression a set of three 150 mm edge cubic samples, cast along with the column and cured under the same environmental conditions. Once the average value of the cubic strength Rcm was evaluated, the corresponding cylindrical one was obtained from the relationship fcm = 0.83Rcm. Furthermore, some members were reinforced with smooth longitudinal steel rebars, commonly used in the past; the remaining ones were realized by using deformed rebars. The longitudinal reinforcement consisted of 14 steel rebars having a 14 mm diameter; according to old italian code provisions, the rebars lap splice length at the column-base joint was always equal to 40 bar diameters (600 mm). In the case of smooth rebars, the end-anchorage consisted of 50-mm radii hooks as typically adopted in the past; a clear concrete cover of about 30 mm was considered at the both sides of the cross section. The transverse reinforcement consisted of 8 mm diameter steel stirrups, 200 mm spaced and closed with 90-degree hooks at both ends (see Fig. 1). The average values of the mechanical properties of smooth and deformed steel rebars, as obtained by tensile tests, are shown in Table 1, where fsy and esy indicate the strength and strain at yielding; esh is the strain at the onset of the steel hardening; fsu and esu are the ultimate strength and the corresponding strain. The systems used for strengthening the test specimens are shown in Fig. 2a and b. The type ‘‘C’’ system (Fig. 2a) consisted of a passive confinement realized by wrapping members with two or four unidirectional CFRP layers. In particular, starting from the column base, the first portion of the member (of about 700 mm) was continuously confined, while the remaining part was strengthened by means of 150 mm spaced strips, each having a width of 100 mm. In order to prevent stress concentrations – which may cause the premature failure of the FRP system – the corners of each column were rounded to a radius of approximately 30 mm before applying the FRP jacket. As provided by the suppliers, the employed CFRP plies were characterized by an elastic modulus of 390 GPa, a tensile strength of 3000 MPa and an ultimate strain equal to 0.80%; the thickness of the single layer was about 0.22 mm. The type ‘‘A1’’ system (Fig. 2b), instead, entailed the use of longitudinal cold bent steel profiles along the member corners before applying the external wrapping made of four CFRP layers; they were four 80  80  6 equal leg angles made of structural steel Fe360. The angles were always glued to the concrete substrate by means of an epoxy adhesive and were anchored to the concrete stub by means of steel connectors; they were properly manufactured in order to have a rounded corner with a radius of about 30 mm. After placing the angles and before realizing the FRP jacket, a 6 mm layer of epoxy mortar was applied on each column side, in order to fill the gap between the profiles (see the detail in Fig. 2b). Fig. 3 shows the anchorage system adopted to connect each longitudinal angle to the concrete stub; in particular, it consisted of a 100  100  15 L-shaped steel base link, 80 mm wide, welded to the longitudinal angle and restrained to the stub by means of a 16 mm threaded rod.

2.2. Test setup and instrumentation The test set up is shown in Fig. 4; it is very similar to that already used for square columns, although slightly modified for taking into account of higher lateral loads expected during these tests.

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(a)

(b)

Fig. 1. Schematic of concrete columns reinforced using smooth (a) and deformed (b) steel rebars.

Loads, strains, displacements and crack widths were measured during the tests; in particular:

Table 1 Mechanical properties of the steel rebars. Type

fsy (MPa)

esy (%)

esh (%)

fsu (MPa)

esu (%)

Smooth Deformed

346 556

0.165 0.265

3.68 3.97

498 655

23.80 16.73

Columns were mounted vertically and tested in displacement control under combined axial and lateral loads. They were restrained to the lab’s floor by means of a steel system which consisted of: (a) two transverse beams placed on the RC foundation and fixed to the floor by means of four high strength thread rods that were properly pre-tensioned in order to avoid any stub rotation; and (b) two stiff steel plates fixed onto the ground and placed orthogonally to the load direction at the bottom of the stub in order to prevent any horizontal movement. The axial load (N) was applied before the horizontal one by pre-tensioning a pair of 40 mm diameter high strength steel rods with a 2000 kN MOOG hydraulic actuator: this actuator, placed at the top of the column, kept the axial load constant during each test. In particular, two values of the normalized compression load ‘‘m’’ – respectively equal to 0.14 and 0.25 – were considered; the dimensionless axial load ‘‘m’’ is given by:

m¼

N fcm  Ag

ð1Þ

where Ag is the area of column cross-section. The horizontal action, instead, was cyclically applied by using a 500 kN MTS hydraulic actuator, mounted at 2050 mm from the column base and fixed to a reaction steel frame. The time history of the horizontal displacement is shown in Fig. 5. In particular, an increment of the imposed horizontal displacement every three cycles was considered in order to evaluate the strength and stiffness degradation at repeated lateral load reversals. After initial cycles at 1, 2, 4 and 6 mm, the displacement amplitude was given as fraction of the estimated tip yield displacement of the column, Dy (15 mm); two different displacement rates were considered during the tests: 0.1 mm/s before the achievement of Dy and 1 mm/s after Dy. Tests were stopped well beyond a predetermined threshold corresponding to the 10% strength degradation evaluated on the monotonic envelope of the load–displacement curves: this threshold represents a ‘‘conventional collapse’’ to which reference will be made in the following for the analysis of the experimental data.

(a) horizontal and vertical strains were monitored using several strain gauges placed on the column at about 100 mm from the stub interface; (b) longitudinal strains of the steel angles were recorded by strain gauges placed on each profile; (c) LVDTs were used to measure potential rigid stub displacements; (d) two potentiometers were used to measure lateral displacements at 2050 mm and at 2500 mm from the column base (i.e. where the lateral load is applied and at the top of the column); and (e) vertical displacements and crack widths at the column-stub interface were monitored by two LVDTs for each side placed at about 100 mm from the base; one of them had the pin located at 30 mm from the column base: in this way, the difference between the readings of two LVDTs allowed estimating the crack opening at the base and the slip of the rebars.

3. Strength and ductility of tested specimens 3.1. Cyclic behavior Table 2 summarizes the main data and results of the nine performed tests. In the table each specimen is identified by a label which provides the column number (from C25 to C33); the longitudinal steel reinforcement type (D stands for deformed rebars; S for smooth rebars); the strengthening system (types ‘‘C’’ or ‘‘A1’’), if any. In addition, for each test the table reports: the number of the FRP layers (2 or 4); the average value of the cylinder compressive strength of concrete (fcm); the applied axial load (N) and its corresponding normalized value (m); the peak lateral strength in the two  direction of loading ðF þ max and F max Þ and the corresponding displacement (D+ and D); the maximum displacement of the column (Dmax) measured at the conventional collapse (i.e. at the achievement of 10% strength degradation evaluated on the horizontal force–displacement F–D curve); the observed failure modes.

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100 100 100 100 100

150

CFRP jacket

700

150

1950

2500

150

150

150

Type "C"

700

(a)

100 100 100 100

cold bent steel profile 80x80x6

100

150

150

150

150

150

Type "A1"

680

mortar

CFRP jacket 700

(b) Fig. 2. FRP strengthening systems.

From Table 2 it is observed that, despite the common belief on the reduced efficacy of the FRP confinement in the case of high aspect ratio, the performed tests have highlighted a non-negligible increase of ductility of the confined members. The positive effect of the FRP confinement on the column performances is more evident in the case of higher axial load level (compare results of tests on column C31 and column C33). Furthermore, as already observed for square members [14], a significant increase of flexural strength is evidenced when the FRP confinement is associated with the use of steel angles connected to the foundation. Fig. 6 shows the lateral force-tip displacement cyclic curves relative to some specimens reinforced with smooth (Fig. 6a and b) or deformed steel rebars (Fig. 6c and d), all subjected to the highest value of axial load (m = 0.25). In particular, Fig. 6a and b compare the behavior of the unstrengthened column reinforced with smooth rebars (test C29) – i.e. the ‘‘reference’’ column – with those of the corresponding

members confined with FRP and strengthened with the ‘‘type A1’’ system (test C27 and C28, respectively). Fig. 6c and d, instead, show similar comparisons obtained in the case of columns reinforced with deformed rebars (i.e. specimens C29, C31 and C32). Regardless of the type of steel reinforcement, the curves relative to FRP confined members are characterized by lower strength degradations when compared with those of the unconfined members. The latter, once the maximum flexural strength is attained, are not able to undergo significant deflection without a considerable reduction of strength; therefore, the conventional collapse is achieved for low values of displacement. In the case of columns strengthened with both FRP wrapping and steel angles (system type ‘‘A1’’), the flexural strength degradation is affected by the occurrence of the steel connector pullout from the concrete stub. At that time, in fact, the contribution of the anchored angles abruptly becomes negligible.

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Ø18 angle 80x80x6 weld

15

80

weld

41 100

weld

15

6

100

weld

Fig. 3. Details of the anchorage in the type A1 system.

Fig. 4. Test setup.

Fig. 5. Displacement history.

As already observed for square members, the cyclic behavior of columns reinforced with smooth rebars is always characterized by a significant ‘‘pinching effect’’, so that a lower energy dissipation capacity has to be attributed to these members.

This appears clearly from Fig. 7, where some hysteretic cycles at equal displacement amplitude, recorded in case of FRP confined columns reinforced by using smooth (test C27) and deformed steel rebars (test C31), are shown. Finally, Fig. 8a and b depict the experimental force–displacement envelopes obtained for members reinforced with smooth or deformed steel rebars, respectively. The comparisons allow to confirm the beneficial effects deriving from the selected strengthened techniques. As expected, the behavior of columns strengthened using the ‘‘type A1’’ system was always characterized by initial stiffness higher than that exhibited by the other members. The use of only FRP confinement, instead, did not affect the behavior of columns in terms of stiffness. It is worth noting that the lower stiffness exhibited by the specimen C33 with respect to the other members is due to some technical issues raised when testing such column; these problems

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Table 2 Test specimens: main data and test results. Test

# FRP fcm layers (MPa)

N (kN) m ()

Fþ max (kN)

D+ (mm)

F max (kN)

D (mm)

C25-S C26-S-C

– 2

12.1 16.3

635 856

0.25 0.25

161.67 229.31

29.43 44.31

158.08 218.63

22.25 43.30

43.46 59.45

C27-S-C C28-S-A1

4 4

13.1 13.8

688 725

0.25 0.25

202.09 302.04

52.29 89.76

215.49 303.90

59.29 89.79

73.59 101.27

C29-D

–

11.7

614

0.25

177.62

22.32

180.54

22.33

34.03

C30-D-C C31-D-C C32-D-A1

2 4 4

22.6 15.0 11.7

1100 787 614

0.25 0.25 0.25

317.57 293.27 319.33

89.57 89.32 66.75

317.95 280.56 320.90

74.65 66.51 66.83

94.91 84.05 69.91

C33-D-C

4

12.6

370

0.14

204.52

47.81

171.45

71.79

64.98

led to prematurely stop the loading process and restart the test again with the member already cracked. 3.2. Dimensionless envelope curves The performances of the tested members can be better compared through the normalized bending moment values ‘‘l’’, which allow to by-pass the dependence of test results on the concrete strength level, thus providing a more immediate comparison of results in terms of flexural strength. The normalized bending moment is given by:

l¼

F  Ls B  H2  fcm

¼

M B  H2  fcm

ð2Þ

where B and H are the width and the depth of the column cross section, respectively; F is the horizontal force applied by the MTS actuator, while Ls is the shear span of the column.

Dmax (mm)

Failure mode Concrete crushing, stirrups opening, and rebars buckling FRP fracture, spalling of concrete cover, stirrups opening, and rebars buckling FRP fracture, and limited phenomenon of concrete crushing Progressive failure of the L-shaped base link with a partial pullout of the steel rods Spalling of concrete cover, rebars buckling, stirrups opening, concrete crushing FRP fracture, rebar failure FRP fracture, rebar failure Complete pullout of the threaded rods with a severe damage of the base link FRP fracture, and concrete spalling

Therefore, in order to better investigate the benefits obtained through the considered strengthening techniques, the experimen tal results already shown in Table 2 ðF þ max ; F max and Dmax Þ are further reported in Table 3, but in a dimensionless form. In particular, Table 3 reports, for each performed test, the peak values of the normalized bending moment (l) in the two directions of  loading ðlþ max and lmax Þ, and the lateral drift ratio ‘‘dmax’’, obtained by dividing the tip deflection of the column at the conventional collapse (Dmax) by the shear span (Ls = 2050 mm). As shown, in the case of members reinforced with smooth rebars, an FRP confinement system made of just two CFRP layers is sufficient to enhance the deformation capacity of rectangular specimens: the column C26 showed an increase of drift ratio over the unstrengthened column C25 (the control specimen) equal to about 37%. In the case of the specimen C27, instead, the use of a four layers system also allows to increase the flexural strength over the member C26 of about 21%; in this case, the benefit in terms of duc-

Fig. 6. Load–displacement hysteretic curves: columns with smooth (a and b) and deformed steel rebars (c and d).

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R. Realfonzo, A. Napoli / Construction and Building Materials 37 (2012) 606–620 Table 3 Normalized values of test results. TEST

# FRP layers

fcm (MPa)

N (kN)

m

lþmax

lmax

dmax (%)

C25-S C26-S-C C27-S-C C28-S-A1

– 2 4 4

12.1 16.3 13.1 13.8

635 856 688 725

0.25 0.25 0.25 0.25

0.19 0.20 0.22 0.31

0.18 0.19 0.23 0.31

2.12 2.90 3.59 4.94

C29-D C30-D-C C31-D-C C32-D-A1 C33-D-C

– 2 4 4 4

11.7 22.6 15.0 11.7 12.6

614 1100 787 614 370

0.25 0.25 0.25 0.25 0.14

0.21 0.20 0.27 0.38 0.23

0.21 0.20 0.26 0.38 0.19

1.66 4.63 4.10 3.41 3.17

Fig. 7. Shape of hysteresis loops.

tility is even more evident, with a percentage increase going up to 69%. Similar considerations can be drawn for members reinforced with deformed rebars. In particular, specimens C30 and C31 showed increases of strength and/or ductility over the unconfined member C29: these increments are greater than those obtained from the corresponding members reinforced with smooth rebars. Finally, regardless of the type of longitudinal steel rebars, columns strengthened with the ‘‘type A1’’ system (i.e. columns C28 and C32), always provided the greatest enhancement of flexural strength; conversely, as mentioned earlier, the increase in terms of ductility is strongly influenced by the performance of the steel connection between the angles and the concrete stub, which affects the achievement of the conventional collapse. In fact, in the case of the specimens C32, the collapse occurred in correspondence of a drift ratio value lower than that obtained by the companion member strengthened only with four CFRP layers (test C27), because of a premature occurrence of the connectors pullout from the column stub. Fig. 9a and b shows the experimental l–d envelopes plotted for members reinforced with smooth or deformed steel rebars, respectively. The l–d comparisons clearly confirm the benefits produced by the FRP confinement and the significant increase of strength exhibited by members strengthened with the system ‘‘type A1’’. It is highlighted that the behavior of these members, once the collapse of anchor system is attained, becomes very similar to that of FRP confined members (compare tests C32 and C31). 3.3. FRP confinement efficiency The efficiency of the adopted retrofitting techniques can be better investigated by computing the ratios ‘‘Id’’, between the maxi-

mum drift of each strengthened column and that of the corresponding control sample:

Id ¼

dstrengthened max

ð3Þ

dunstrengthened max

and the ratios ‘‘Il’’ between the flexural strengths:

Il ¼

lstrengthened av lunstrengthened av

ð4Þ

being lav the average value of the peak normalized bending moment l evaluated in pull and push direction of loading  ðlþ max and lmax Þ:

lav ¼

lþmax þ lmax 2

¼

ðF þmax þ F max Þ=2  Ls B  H2  fcm

ð5Þ

The values of these indices are reported in Table 4. In order to allow an overall evaluation of the FRP confinement efficiency by varying the shape of the cross-section, the table also provides the results of tests performed on square columns; details about tests on these square members and comments related to the results in the table can be found in [14]. Furthermore, it is mentioned that, in the case of rectangular members, strengthening systems made of GFRP confinement or those having CFRP wrapping and steel profiles unconnected to the foundation (systems ‘‘type A2’’) have not been investigated. It is worth noting that the values of the indices Id and Il for the test C33-D-C are not shown in the table as no companion unstrengthened specimen has been tested under m = 0.14. In the case of rectangular members, data shown in the table highlight that:

Fig. 8. Load–displacement envelopes: columns with smooth (a) and deformed steel rebars (b).

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Fig. 9. Dimensionless l–d envelopes: columns with smooth (a) and deformed steel rebars (b).

Table 4 Drift and strength increase. Cross section

m

TEST

FRP layers

Il

Square

0.14

C4-S-G C1-S-G C10-S-C C13-S-C C2-S-A1 C11-S-A1 C6-S-A2

2 4 2 2 2 2 2

1.05 1.02 0.96 0.83

C7-D-C C8-D-C C14-D-C C15-D-A1

2 2 2 2

1.19 1.16 0.80

C17-S-C C19-S-C C20-S-A1 C12-S-A2

2 4 2 2

1.19 1.25

C22-D-C C23-D-C C24-D-A1

2 4 2

1.15 1.33

C26-S-C C27-S-C C28-S-A1

2 4 4

1.04 1.21

C30-D-C C31-D-C C32-D-A1

2 4 4

0.95 1.26

Type ‘‘C’’

0.40

Rectangular

0.25

Id Type ‘‘A1’’

Type ‘‘C’’

Type ‘‘A1’’

1.99 2.03 1.35 1.60 1.45 1.36

1.84 2.15 1.99 1.83 1.84

2.33

1.42 1.70 1.93

2.55

2.10 2.00 2.95

1.81

1.52 1.37 1.69

1.66

2.33 2.79 2.47

1.81

Fig. 10. Unconfined column with smooth steel rebars (specimen C25-S).

2.05

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Fig. 11. FRP confined columns with smooth steel rebars (C26-S-C & C27-S-C).

Fig. 12. Unconfined column with deformed steel rebars (C29-D).

– the supposed reduction of the confinement effectiveness related to the high aspect ratio is truly negligible since benefits in terms of ductility attained by wrapping rectangular columns with CFRP are comparable with those obtained by confining square members; – although limited, a strength increase due to the FRP jacket (about 20%) is found when using four FRP layers; and – with respect to the reference (unstrengthened) members, benefits similar to those obtained for square columns can be observed for specimens strengthened by combining the use of steel angles and FRP jacket (type A1 systems): improvements in terms of strength up to 81% and of ductility up to 133% have been obtained. 4. Failure mode During tests, in the case of both unconfined and FRP confined specimens, columns reinforced with smooth steel rebars exhibited cracking phenomena and significant damages concentrated in the first 500 mm from the column base (see Figs. 10 and 11). In particular, due to the low bond strength between concrete and smooth rebars, a large crack occurred at column-stub interface whose width progressively increased during the test; this crack was even more dominant in presence of FRP confinement. Members reinforced with deformed rebars, instead, showed a considerable level of damage in a larger portion (about 800 mm from the column base) when unconfined (Fig. 12). The presence of a FRP jacket inhibited the crack opening (Figs. 13 and 14); concrete cracks developed and propagated only in those member

regions characterized by wrapping discontinuities (i.e. the column base and the portion at about 600–700 mm from it). Regardless of the type of reinforcement (smooth or deformed), the presence of a FRP jacket has always allowed to delay the crushing of concrete cover and the subsequent buckling of longitudinal rebars. Fig. 15 shows the damage exhibited by specimens strengthened using FRP jacket and steel profiles (columns C28, reinforced with smooth rebars, and C32, having deformed rebars). For these members, only a large crack at the column base was observed during the test, whose width reached about 25 mm; the conventional collapse was always attained at the failure of the connection between the steel profile and the concrete foundation, which in turn was caused by the pullout of the connectors from the concrete stub. The following sections better examine the crack propagation and the failure mode characterizing all the different typologies of tested specimens. 4.1. Unconfined and CFRP confined columns reinforced with smooth steel rebars Figs. 10 and 11 show the damage level of tested columns as observed at different steps of load, up to the collapse. In particular, in the case of the unconfined member (test C25-S), a flexural crack first occurred at column-stub interface at a tip displacement amplitude of 15 mm (Fig. 10a). Then, the damage rapidly involved the upper portion of the column, where some sub horizontal cracks opened at spacing of about 200 mm, i.e. at steel

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615

Fig. 13. FRP confined columns with deformed steel rebars with m = 25% (C30-D-C & C31-D-C).

Fig. 14. FRP confined column with deformed steel rebars with m = 14% (C33-D-C).

stirrups spacing. These cracks progressively developed in diagonal directions and were accompanied by vertical cracks due to the incipient buckling of the rebars in compression (Fig. 10b). Fig. 10c shows the significant width of these vertical cracks with subsequent occurrence of concrete spalling at the achievement of the ‘‘conventional collapse’’ (corresponding to a column drift of about 60 mm). Crushing of concrete, buckling of rebars and stirrups opening are clearly evidenced by Fig. 10d; such damage phenomena involved a length of about 500 mm measured from the column base. As for the unconfined column, in the case of FRP confined ones (tests C26-S-C and C27-S-C) a first flexural crack opened at the base at a displacement amplitude of 15 mm (Fig. 11a), whose width progressively increased during the test. For a drift level of 45 mm both abovementioned columns evidenced an initial fracture of the FRP jacket, at about 500 mm from the base, i.e. at the end of rebars lap-splicing. Fig. 11b shows the damage in the overlapping zone of the rebars towards the end of the test. When using two carbon layers (column C26), the FRP jacket failure rapidly propagated up to the column base; this abrupt failure of the FRP system gave rise to the following chain damage phenomena: spalling of the concrete cover, then stirrup opening and, finally, rebar buckling (Fig. 11c). The damaged region was located in the first 500 mm from the concrete stub. The use of four FRP layers (specimen C27) led to a more gradual propagation of the FRP fracture, so that at the end of the test, concrete failure was only observed at the column base (for a smaller portion of about 100 mm). Fig. 11d refers to a picture taken at

the end of this test once the FRP wrapping was removed: it is clearly evidenced that the portion of the column hidden by the FRP wrapped is mostly characterized by absence of damage, except for the base section where a limited phenomenon of concrete crushing is observed. 4.2. Unconfined and CFRP confined columns reinforced with deformed steel rebars The unconfined member (test C29-D) exhibited a significant level of damage that involved a length of about 800 mm from the column base. The first signs of damage occurred at a low displacement value (15 mm) and were characterized by the crack opening at column-foundation interface and by the development of diagonal cracks at 600–800 mm from the column base (Fig. 12a). A further crack opened at about 600 mm from the base (i.e. at the end of the rebars lap splice length) and it was followed by the opening of other cracks distributed in correspondence to the steel stirrups (i.e. 200 mm spacing). Increasing the lateral displacement, a progressive development of vertical cracks was observed, due to the incipient buckling of the compressed rebars (Fig. 12b). This phenomenon led to the concrete spalling first at the column base (Fig. 12c) and then at higher locations (at 700 mm), up causing the collapse of the member with the stirrups opening and the longitudinal rebars buckling (Fig. 12d). In presence of a FRP jacket (tests C30-D-C, C31-D-C and C33-DC), instead, the crack pattern was characterized by the usual crack at the column base section and by further cracks concentrated in a

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Fig. 15. Columns strengthened with FRP and steel devices (specimens C28-S-A1 & C31-D-A1).

small portion at about 600–700 mm from the column base, i.e. close to the end of the continuous FRP wrapping (Figs. 13 and 14). In the case of tests C30 and C31, performed under the higher value of the axial load (m = 0.25), the damage was mostly concentrated at the column base. In particular, for an imposed lateral displacement of about 45 mm the width of the flexural crack at the column base became significant. Subsequently, for a drift of about 90 mm, the FRP jacket failed close to the column-stub interface (Fig. 13a). Fig. 13b, taken at the end of the test C30 after removing the FRP wrap, shows the damage concentration in the first 100–150 mm from the column base. By focusing on this area slight signs of buckling of the compressed rebars were observed while some longitudinal rebars experienced a tensile failure probably due to fatigue phenomenon (Fig. 13c). In the case of column C33, tested under the lower axial load (m = 0.14), the first crack was always experienced at the column base, although under a minor drift value of about 20 mm. As for the unconfined members, this crack did not significantly enlarge during the test. As shown in Fig. 14a, increasing the drift the most damaged zone (only 150 mm wide) shifted at about 600–700 mm from the column base, where concrete spalling was observed at the end of test (Fig. 14b). Fig. 14c, taken at the end of the test C33, evidences how the portion of the column ‘‘hidden’’ by the FRP wrapping was characterized by absence of damage; in fact, slight signs of concrete crushing were observed at the column base, and buckling phenomena of steel rebars were not experienced.

4.3. Columns strengthened with FRP and steel angles Fig. 15 shows the damage observed during the tests C28-S-A1 and C32-D-A1 up to the collapse. As mentioned earlier, regardless of the steel reinforcement type, only a large flexural crack was observed during the test. It opened at the column-foundation interface for a drift of 30 mm and progressively enlarged up to reach a width of about 25 mm under an imposed displacement of 90 mm (Fig. 15a). In both tests, the cyclic behavior was influenced by the pullout of the connectors from the concrete stub that was recorded soon after the opening of the abovementioned flexural crack. Nevertheless, after an initial similar stage, the base connection behaved differently for the two specimens. In the case of test C32 the pullout of the threaded rod became early relevant (Fig. 15b), and for an imposed drift of about 60 mm the contribution of the angles to the column flexural strength dropped down thus causing the achievement of the conventional collapse. At this

displacement level the L-shaped base link exhibited significant deformations. The abrupt loss of strength due to the complete pullout of the rods is clearly shown in both cyclic curve (Fig. 6) and monotonic envelope (Fig. 8). In the case of column C28, instead, a more gradual pullout of the threaded rod was observed and the contribute of the angles to the member flexural behavior was assured up to a displacement of about 100 mm; after this threshold one of the L-shaped base links progressively failed. The angle rupture occurred on the corner between the two edges as shown in Fig. 15c. Finally, it is worth highlighting that, in both tests, no FRP fracture was observed. 5. Chord rotations vs drift ratio curves As highlighted earlier, the longitudinal rebars type and the axial load level have both a strong influence on the evolution of the cracks and consequently on the column cyclic responses. This is the reason why it is useful to focus the attention on local performances of the columns such as the column base rotation. For this purpose, the plots in Fig. 16 depict, for some FRP confined columns, the relationships between the rotation (h) at the base of the specimens (local deformation parameter) and the drift ratio d (global deformation parameter). Such rotation was obtained by the data recorded with the LVDTs located at the bottom of the column, on the two sides perpendicular to the loading direction. In particular, Fig. 16a shows the dependence of the response on the axial load level for two columns reinforced by deformed steel rebars. It can be observed that under m = 0.14 (test C33) the rotation at the base nearly coincides with the drift ratio. This implies that the rotation capacity of the member is slightly influenced by the flexural contribution, while the deformation component due to the rigid rotation at the base is dominant. Similar considerations were already drawn from test results on square columns performed under m = 0.14 [14], where, mainly when using smooth rebars, the rigid rotation is generated by the slippage and elongation of such rebars at the column-foundation interface where a significant cracking is observed. Similar results were also found by other Authors [8,15]. Under m = 0.25 (test C31), instead, the slope of the h–d curves significantly reduces as the flexural contribution to the rotation capacity becomes relevant. The influence of the longitudinal steel reinforcement type is instead shown in Fig. 16b; the figure depicts the cyclic response of columns reinforced with smooth (test C26) and deformed rebars (test C31) and subjected to the same axial load level. From the figure it can be observed that, always due to the lower bond characterizing the smooth rebars, the contribution due to the

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617

Fig. 16. Comparison of h–d curves: influence of the axial load (a) and steel reinforcement type (b).

rigid rotation exhibited by column C26 is much greater than that observed for column C31. 6. Stiffness degradation and energy dissipation Based on the experimental results, it was possible to evaluate the mean value of stiffness for the i-th cycle by using the following ratio [16]:

k¼

jF þmax;i j þ jF max;i j jsþmax;i j þ jsmax;i j

ð6Þ

The stiffness of each displacement cycle k was then normalized with respect to that of the first cycle kI, thus providing a measure of the stiffness degradation.

The relationships between k/kI and drift ratio are plotted in Fig. 17a and b for rectangular specimens reinforced with smooth and deformed rebars, respectively, all subjected to m = 0.25. As it can be first observed, the relationship between the stiffness degradation and the drift ratio is practically independent on the presence of the FRP confinement system; analogous considerations can be found in [17]. In addition, columns reinforced with smooth steel rebars exhibit progressive stiffness degradation only slightly higher than the counterparts reinforced with deformed ones. It is highlighted that the plot relative to the test C31 cannot be considered in the comparative analysis; in fact, such specimen shows an unexpected stiffness degradation because it was been subjected to some previous damage cycles.

Fig. 17. Stiffness degradation vs drift ratio: columns with smooth (a) and deformed steel rebars (b).

Fig. 18. Dissipated energy vs drift ratio: columns with smooth (a) and deformed steel rebars (b).

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Fig. 19. FRP strains.

Furthermore, by considering constant deformation cycles, the comparisons show that, regardless of the steel reinforcement type, columns strengthened with FRP and angles (C28 and C32) have always exhibited the lowest stiffness degradation. Finally, Fig. 18a and b depict the relationship between the cumulative dissipated energy (E) and the imposed drift ratio, for columns reinforced with smooth and deformed rebars, respectively. This energy parameter was calculated from the area under

the lateral load-drift response enclosed within one complete cycle up to the achievement of the conventional collapse for each specimen. The test C33, the only performed under m = 0.14, has been omitted in this comparison. As observed from tests on square columns [14], regardless of the steel reinforcement type, the FRP confinement produces a significant increase of the cumulative dissipated energy; however, in comparison with unconfined specimens, there is no remarkable

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R. Realfonzo, A. Napoli / Construction and Building Materials 37 (2012) 606–620 Table 5 FRP strain values. e (D+) (‰)

e ðDmax Þ (‰)

emax (‰)

0.15 0.79 0.00 0.87 1.04 0.39

0.20 0.91 0.08 0.96 1.11 0.49

0.01 0.00 0.08 0.02 0.01 0.03

0.51 2.69 0.04 0.22 1.08 0.32

0.05 1.05 0.24 0.20 0.48 0.29

1.14 – 0.83 0.19 1.08 0.32

2.14 0.02 1.75 0.03 0.01 0.01

0.13 0.75 0.69 1.00 0.13 0.44 0.31 0.06

0.25 0.75 0.87 1.22 0.42 0.96 0.43 0.14

0.07 0.06 0.25 0.67 0.38 0.96 0.24 0.13

0.06 0.13 0.06 0.88 0.38 0.81 0.13 0.13

0.38 0.14 0.05 0.46 0.17 0.08 0.09 0.14

0.69 1.10 0.65 0.15 0.50 0.05

– 0.10 1.28 0.15 0.10 0.21

0.69 1.10 0.65 0.68 1.61 0.60

0.07 0.50 0.05 0.56 0.72 0.53

0.70 0.50 0.38 0.38 0.90 0.08

0.70 0.10 1.28 0.38 0.10 0.21

0.008 0.024 0.025 0.005 0.027 0.002

0.15 0.98 0.10 1.10 0.02 0.84

– 0.10 1.28 0.15 0.10 0.21

0.23 0.98 0.31 0.73 2.32 0.46

0.36 0.39 0.17 0.62 1.83 0.45

0.70 0.50 0.38 0.38 0.90 0.08

0.70 0.10 1.28 1.10 0.10 0.84

0.014 0.020 0.022 0.032 0.032 0.021 0.019 0.012

0.06 0.36 0.50 0.24 0.13 0.33 0.06 0.36

0.08 0.35 0.40 0.28 0.25 0.38 0.05 0.38

0.09 0.42 0.50 0.18 0.23 0.56 0.36 0.17

0.27 0.26 0.02 0.71 0.39 0.56 0.35 0.22

0.25 0.25 0.08 0.53 0.65 0.13 0.18 0.45

0.44 0.04 0.12 0.71 0.71 0.01 0.01 0.46

e ðDþmax Þ (‰)

eþmax (‰)

Test

Strain gauge

e (D = 0) (‰)

C26-S-C

S1 S2 S3 S4 S5 S6

0.035 0.079 0.044 0.006 0.009 0.008

0.68 2.07 0.42 0.18 0.56 0.10

0.93 1.84 0.36 0.19 0.70 0.09

0.93 2.29 0.44 0.97 1.13 0.49

C27-S-C

S1 S2 S3 S4 S5 S6

0.032 0.028 0.010 0.007 0.016 0.010

0.51 2.35 0.27 0.01 0.16 0.03

2.14 – 1.64 0.03 0.37 0.01

C28-S-A1

S1 S2 S3 S4 S5 S6 S7 S8

0.008 0.016 0.050 0.062 0.030 0.036 0.017 0.007

0.01 0.69 0.81 1.22 0.14 0.35 0.28 0.11

C30-D-C

S1 S2 S3 S4 S5 S6

0.019 0.019 0.028 0.023 0.015 0.022

C31-D-C

S1 S2 S3 S4 S5 S6

C32-D-A1

S1 S2 S3 S4 S5 S6 S7 S8

improvement of the energy dissipated per cycle. In the case of columns reinforced with deformed rebars, doubling the number of the FRP layers (from two to four) produced a small increase of the energy performance (compare test C30 and C31) that, however, was not corroborated by the corresponding comparison from specimens reinforced with smooth rebars (compare tests C26 and C27). Finally, a greater per cycle-energy is dissipated by the columns strengthened with the ‘‘type A1’’ system with respect to the unconfined and FRP confined specimens. However, a reduction of the total dissipated energy is computed for the test C32 because of an early achievement of the conventional collapse.

7. FRP strain As already mentioned, during tests measurements of the hoop strain of the FRP jacket were recorded by several strain gauges (six gauges in the case of ‘‘type C’’ columns and eight for the ‘‘type A1’’ ones); according to the schematics of Fig. 19, the gauges were arranged on the FRP jacket at a distance of about 100 mm from the column base. Fig. 19 also shows for all strengthened members tested under m = 25% (hence except for the test C33) the envelopes of the FRP transverse strains (e) obtained under both positive and negative directions of the imposed displacement D. In particular, the strain

e (D) (‰)

diagrams refer to the maximum values recorded at the first of the three cycles performed at each imposed displacement level, up to the conventional collapse; negative and positive values of e indicate compressive and tensile strains of the FRP jacket, respectively. Under D = 0, i.e. at the origin of the e–D axes, the plots report the small values of the FRP strains exhibited after the application of pure axial load on the columns. The intersections between each strain envelope curve and the vertical dotted lines, plotted at D+ and D, allow to identify the values of the FRP strains corresponding to the achievement of the peak lateral load in both push and pull directions. In the case of columns strengthened by only CFRP confinement (Fig. 19a through Fig. 19d), the plots show similar trend independently on the steel reinforcement type (smooth or deformed). The maximum strain values in tension were always recorded by strain gauges S2 or S5, i.e. those placed at the center of the two opposite sides of the member perpendicular to the load application. These gauges provided strain values significantly higher than the others, thus anticipating the location of the first FRP fracture at the column base (if any). Such maximum tensile values were not measured in correspondence of the achievement of peak lateral load and never attained the ultimate FRP strain value indicated by the supplier (ef,u = 8‰). In particular, the highest value of the FRP strain, equal to 2.69‰, was measured by gauge S2 during the test C27, which is only 1/3 of ef,u.

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The strain measured in tension by gauges S2 and S5 was always higher for columns confined by four CFRP plies (see tests C27 and C31). Conversely, the maximum strain values measured by the other gauges (i.e. S1, S3, S4, and S6) were generally slightly higher in the case of members confined with two CFRP layers (see tests C26 and C30). In the case of columns strengthened with the system ‘‘type A1’’ (Fig. 19e and f), the use of steel profiles seems to reduce the hoop strains exhibited by the FRP jacket; as it can be observed for the test C32, the maximum tensile strains do not exceed the value of 0.56‰, i.e. corresponding to 1/14 of ef,u. However, it has to be highlighted that the gauges were always placed close to the steel angles and not at the center of the two opposite sides perpendicular to the load application. In this way, the highest strain values of the FRP jacket might not have been recorded. The considerations above reported can be finally summarized through Table 5, listing the main results on the FRP strains obtained from the gauges. In particular, for each test the table reports the strain values measured in both pull (+) and push (–) directions under the following displacements values: – D = 0, i.e. at the beginning of the test soon after the application of the axial load; – D+ and D–, i.e. at the achievement of the peak lateral load; and  – Dþ max and Dmax , corresponding to the achievement of the conventional collapse. The table also reports in italic the peak strain values measured  in tension ðeþ max Þ and compression ðemax Þ by each strain gauge; the highest strain values recorded in tension for each test are marked in bold. Such information will be of course valuable for future developments of the present research aimed at better examining the efficiency of the FRP systems for strengthening RC structural members. 8. Concluding remarks In this paper, results from cyclic tests performed on full scale RC columns, having a 300  700 mm cross section and strengthened with FRP systems, have been discussed. Despite the high side ratio (more than 2.3), it has been shown that the FRP confinement leads to a significant ductility enhancement of the columns; in fact, the presence of a FRP jacket considerably reduces their damage level by inhibiting the development of the crack pattern and delaying the spalling of the concrete cover and the subsequent buckling of the longitudinal rebars. The combined use of FRP confinement and steel angles, instead, allows a significant increase of flexural strength although a reduction of the available ductility has been observed in one of the two performed tests; this loss of ductility was due to the brittle failure of the connection between the steel profile and the concrete foundation which in turn was caused by the pullout of the connectors from the concrete stub.

For what concerns the stiffness degradation and the energy dissipation the experimental results have shown that: the relationship between the stiffness degradation and the drift ratio is practically independent on the presence of the FRP confinement system; also in the case of high aspect ratio columns the FRP confinement produces a significant increase of the cumulative dissipated energy. Finally, it has been observed that the maximum strain values of the FRP jacket never attained the ultimate value indicated by the supplier (ef,u = 8‰); in particular, the highest value of the FRP strain recorded during tests was equal to 2.69‰ which is only 1/3 of ef,u. Acknowledgement The authors gratefully acknowledge ‘‘Interbau srl’’ firm (Milan, Italy) for the financial support of the experimental program. References [1] American Concrete Institute (ACI). Guide for the design and construction of externally bonded FRP systems for strengthening concrete structures. ACI 440.2R-08. Farmington Hills, Mich. [2] Fédération internationale du Béton (fib). Externally bonded FRP reinforcement for RC structures. Bulletin No. 14, technical report. Lausanne, Switzerland. [3] CNR-DT200. Guide for the design and construction of externally bonded FRP systems for strengthening existing structures (materials, RC and PC structures, masonry structures). CNR, National Research Council, Rome, Italy; 2004. [4] Teng JG, Chen JF, Smith ST, Lam L. FRP-strengthened RC structures. UK: Wiley; 2002. [5] Realfonzo R, Napoli A. Concrete confined by FRP systems: confinement efficiency and design strength models. Composites: Part B 2011;42:736–55. [6] Iacobucci RD, Sheikh SA, Bayrak O. Retrofit of square concrete columns with carbon fiber reinforced polymer for seismic resistance. ACI Struct J 2003;100(6):785–94. [7] Harries KA, Ricle JR, Pessiki S, Sause R. Seismic retrofit of lap splices in nonductile square columns using carbon fiber-reinforced jackets. ACI Struct J 2006;103(6):874–84. [8] Harajli MH, Dagher F. Seismic strengthening of bond-critical regions in rectangular reinforced concrete columns using fiber-reinforced polymer wraps. ACI Struct J 2008;105(1):68–77. [9] Bousias SN, Fardis MN, Biskinis D. Retrofit of RC columns with deficient lap splices. Fib symposium ‘‘Keep Concrete Attractive’’, Budapest; 2005. [10] Bournas DA, Lontou PV, Papanicolaou CG, Triantafillou TC. Textile-reinforced mortar versus fiber-reinforced polymer confinement in reinforced concrete columns. ACI Struct J 2007;104(6):740–8. [11] Bousias SN, Triantafillou TC, Fardis MN, Spathis L, O’Regan BA. Fiber-reinforced polymer retrofitting of rectangular reinforced concrete columns with or without corrosion. ACI Struct J 2004;101(4):512–20. [12] Harajli MH. Behavior of gravity load-designed rectangular concrete columns confined with fiber reinforced polymer sheets. J Compos Constr 2005;9(1):4–14. [13] Napoli A, Nunziata B, Realfonzo R. Cyclic behaviour of rectangular reinforced concrete columns strengthened with FRP systems. In: Proceedings of XIII national congress ANIDIS, Bologna; 28 June–2 July 2009. [14] Realfonzo R, Napoli A. Cyclic behavior of RC columns strengthened by FRP and steel devices. J Struct Eng 2009;135(10):1164–76. [15] Fabbrocino G, Verderame GM. Rotational capacity of old type R/C columns. In: Proceedings of fib symposium, Budapest, Hungary; May 23–25, 2005. [16] Mayes RL, Clough RW. State of the art in seismic shear strength of masonry – an evaluation and review. EERC Report 1975, Berkeley, California, USA. [17] Wu Y, Liu T, Wang L. Experimental investigation on seismic retrofitting of square RC columns by carbon FRP sheet confinement combined with transverse short glass FRP bars in bored holes. J Compos Constr 2008;12(1):53–60.