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Hybrid Confinement of Concrete by FRP sheets and Fiber Ropes Under Cyclic Axial Compressive Loading
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Theodoros C. Rousakis1
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Abstract: The study looks into the mechanical behavior of concrete lightly confined by glass
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Fiber Reinforced Polymer (FRP) sheets and by polypropylene Fiber Ropes (PPFRs). PPFRs have
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ultrahigh deformation at failure. The hybrid confining technique suggests applying the fiber rope
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on already cured FRP jackets as external unbonded reinforcement. No impregnation or gluing
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resins are necessary for the FR application. The technique uses mechanical anchorage for the
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ends of the FR. Sixteen confined standard concrete cylinders, in two series and with different
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concrete qualities, were tested under repeated axial compression cycles of increasing
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displacement. They included confinement of one layer of glass FRP and FR confinement in
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different volumetric ratios. The investigation also compares columns confined only by GFRP
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jackets or only by PPFR wrapping. The elaboration evaluates the axial stress versus axial and
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lateral strain behavior of the columns. Proper design of the hybrid confinement utilizes further
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the confining effects of the FRP sheet up to its multiple fracture. Then, FR ensures increased
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axial strain (higher than 5%) and dissipated energy of concrete. The load presents a temporary
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yet controlled loss, followed by an increase for further loading.
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CE Database subject headings: concrete columns; confinement; strengthening; FRP; fiber
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rope; hybrid; cyclic compression; overloads; energy dissipation.
1
Lecturer, Laboratory of Reinforced Concrete, Dept of Civil Engineering, Democritus University of Thrace (DUTh), Vas. Sofias
12, 67100 Xanthi, Greece. E-mail:
[email protected], tel.: 2541079645, fax: 2541079638.
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Introduction
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Confinement of concrete by Fiber Reinforced Polymer (FRP) material is reliable, efficient and
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easy to apply. It utilizes the increased strength, axial strain ductility and energy dissipation of
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concrete when being under an ever-increasing triaxial compression state. Carbon, glass or
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aramide fiber sheets provide the needed lateral restriction to alter concrete from a fragile to a
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rather ductile material. The studies by Rousakis 2001, Karabinis and Rousakis 2002, Tamuzs et
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al. 2006, Rousakis et al. 2008, Rousakis and Karabinis 2008 and 2012, Matthys et al. (2006) and
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Liang et al. (2012), among numerous others present the confining effects of carbon (or glass)
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FRP materials. Carbon sheets are more advantageous than other reinforcing fibers for their time-
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dependent performance and their durability while they have a high modulus of elasticity. Thus,
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they result in a confinement of only a few layers of sheet of small thickness. Fardis and Khalili
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(1981), De Lorenzis and Tepfers (2003), Teng et al. (2009) and more recently Wie and Wu
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(2012) and Rousakis et al. (2012) among others have identified the main design parameters. The
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FRP ultimate lateral pressure and/or its axial rigidity (thickness and modulus of elasticity) are
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used to predict the strength and failure strain of concrete. They both express the FRP lateral
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restriction throughout imposed axial deformations in concrete. Yet, the key role of the axial
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rigidity of the confining reinforcement attracts little attention so far. In that case, the optimum
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confining material (from a mechanical point of view) may be evenly associated with its high
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deformation at failure, rather than the high modulus of elasticity and failure stress. That may be
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the case especially for seismic strengthening, where high concrete strain ductility, and so the
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integrity of concrete core throughout loading, is important. FRPs with high modulus of elasticity
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possess low strain at failure. In fact, the failure in columns confined by conventional FRPs
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usually involves the fracture of the jacket. Therefore, many recent experimental and analytical
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investigations have focused on the better understanding of the crucial FRP strain at fracture (De
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Caso y Basalo et al. 2011, Chen et al. 2011 among others). Anggawidjaja et al. (2006) and Dai et
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al. (2011) examined FRP sheet confinement with materials having low modulus of elasticity and
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thus higher deformability. They used polyethylene naphthalate (PEN) and polyethylene
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terephthalate (PET) instead of carbon or glass fibers. As the modulus of elasticity of the PEN and
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PET materials is lower than that of carbon or glass fibers, they compensated for the needed
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stiffness by providing higher thickness. The high deformability of PEN and PET provided
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concrete with axial strain ductility beyond the requirements of many practical applications.
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After the extensive research and applications on the use of FRP reinforcement to strengthen
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concrete, some recent efforts explore also advanced materials and techniques. One major field in
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concrete confinement concerns the use of inorganic binders instead of the impregnation resins.
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Thus, the fiber reinforcement is in the form of suitable textiles or fiber reinforced sheets.
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Inorganic binders raise the resistance of the confinement strengthening at elevated temperature,
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resist ultraviolet (UV) radiation and extend the use to aggressive environmental conditions. In
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addition, the experimental performance of the various resulting jackets has been satisfactory
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(Triantafillou and Papanicolaou 2005, Bournas et al. 2007, Peled 2007, De Caso y Basalo et al.
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2012).
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Another research field utilizes the unique FRP jacketing characteristics in hybrid strengthening
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schemes. Hu et al. (2011) used external FRP wraps in concrete encased in thin steel tube
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columns to delay or even prevent outward local buckling deformation of the steel tube. Shin and
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Andrawes (2011) combined external FRP wraps and outermost shape memory alloy spirals to
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upgrade the seismic performance of reinforced concrete columns.
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Confining reinforcement in the form of fiber ropes (FR) has attracted little attention so far.
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Aramide or vinylon performs well as external or internal shear reinforcement (Shimomura and
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Phong 2007). They present low sensitivity to local damage of fibers because of handling,
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scratching or stress concentrations in the edges of noncircular sections. They need no
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impregnation resins or binders in external confinement applications. Vinylon has very low
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modulus of elasticity. Polypropylene has an elastic behavior and even lower modulus of
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elasticity. Peled (2007) investigated among other confining reinforcements the behavior of
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concrete cylinders encased into thin polypropylene tubes. The confined concrete showed a
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softening stress-strain response of significant failure strain. In his study, Rousakis (2013)
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investigated Vinylon Fiber Ropes (VFR) and Polypropylene Fiber Ropes (PPFR) as external
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confinement for low strength concrete. Heavily FR confined concrete (non pre-damaged)
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reached remarkable ultimate strains of 13%. The bearing stress was more than 6 times higher
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than that of plain concrete, accompanied by a high energy dissipation of concrete. These values
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were not ‘failure’ ones, as no fracture of the FR reinforcement took place. Because of loading
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machine limitations, a ‘spring – like’ concrete-related ‘failure’ occurred upon full unloading of
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the concrete column. The FR wrapping requires minimum time, as it does not use resin or
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mortar. Resin-free FR ensures rather higher temperature resistance of the retrofit and healthier
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working conditions for workers and engineers. The FR technique could be applied to harsh
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climates and environments of low (or high) temperature and high humidity. PPFRs are reusable
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and recyclable.
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The presented study utilizes the unique, ductile behavior of concrete confined by PPFRs that
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possess ultrahigh tensile deformation at failure. Extreme seismic excitations of structures can
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cause overloading of the critical region of a concrete member. PPFR materials may ensure the
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potential extra demand of the compressive strain ductility of concrete. In columns confined by
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FRP sheets, an abrupt fracture of the confining reinforcement may occur after excessive imposed
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lateral deformations. Then, a precipitous drop of the bearing load takes place. PPFR can preserve
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the integrity of low strength concrete to a remarkable strain level while presenting ever-
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increasing axial load capacity (Rousakis 2013). Thus, this study further examines the confining
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effects of hybrid glass FRP - polypropylene (PP) fiber rope (FR) external confinement. The first
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purpose is to utilize the confinement effects of FRPs for the early stages of inelastic loading. The
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second objective is to avoid the abrupt fragile failure of the concrete member after the fracture of
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the FRP sheet. This fracture may occur in low axial strains, even for multiple layers of FRP
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sheets when the concrete section presents a low confinement effectiveness coefficient.
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Considering the promising results by Shimomura & Phong (2007), the consequent concrete
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confinement investigation is necessary before the extensive experimental and analytical
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elaborations of the effects in reinforced concrete columns under combined flexure-shear-
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compression. The tests include concrete cylinders in two series with different concrete qualities
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under repeated axial compression cycles of increasing displacement. The specimens have only
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one external layer of glass FRP and FR confinement in different volumetric ratios. This study
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presents only the specimens designed to provide an acceptable temporary load drop (less than
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20%).
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Experimental testing
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The study presents twenty-six axial compression tests on standard plain concrete cylinders, with
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150 mm diameter and 300 mm height, constructed by ready–mixed concrete of two different
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compressive strengths. Sixteen specimens had FRP confinement or polypropylene fiber rope
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(PPFR) wrapping or a combination of the two (hybrid confinement). 5
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Materials’ mechanical properties
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Two different compositions of normal strength concrete have been used. The C16 mixture
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(characteristic compressive strength of 16 MPa) had the ratios of water/cement, sand/cement,
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gravel/cement of 0.62, 3.9, 3.0 respectively. The C20 mixture (characteristic compressive
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strength of 20 MPa) had the corresponding ratios of 0.54, 3.2 and 2.44. Three concrete cylinders
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from each batch reached an average 28-day compressive strength of 20.5 MPa for C16 and 25.8
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MPa for C20 mixtures. Four additional concrete cylinders included steel collars, placed at the
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anchorage region of the rope, near the top and bottom base of the specimens (Figure 1a). The
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steel collars were positioned with a relative distance of 255 mm centers (gap of 5 mm from
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cylinder ends). The main use of the collars was to anchor the ends of the rope, while the axial
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displacement transducers were attached on them. These tests took place with the remaining
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sixteen specimens four and a half months after casting. The average concrete strength during the
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tests was 25.1 MPa at a strain of 0.401 % for the C16 batch. The matching stress for the C20
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batch was 33.7 MPa at a strain of 0.389 %. These values were used for further elaborations of the
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externally strengthened columns.
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The glass FRP jacket consisted of a S&P G90/10 type unidirectional sheet with 300 mm width
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(S&P—Sintecno, Scherer 1999). Table 1 presents the mechanical properties of the glass FRP
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sheet. Treatment of any undesirable cavities and pores needed the use of a two-component epoxy
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paste (P103, Sintecno). A primer resin (S2W, Sintecno) and a sheet impregnation resin (S2WV,
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Sintecno) were used to apply the FRP material. More details on the application of the glass FRP
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sheet materials can be found in Rousakis and Karabinis (2012).
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The polypropylene fiber rope (PPFR, product of Thrace Plastic Co. S.A.) is designed with two Z-
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twisted strands. The rope has a tensile modulus of elasticity of 2.0 GPa and a tensile elongation
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at failure of 20.36%. Figure 1c shows the FR while Table 1 includes the mechanical properties.
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Detailing of specimens
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The confined cylinders varied in their concrete strength and in their external strengthening
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scheme. Two columns in each concrete batch were confined by 1 layer of glass FRP (GFRP).
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The sheet wrapping had an overlap of 150 mm to avoid anchorage debonding failure. Three other
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specimens in each batch were wrapped by 5 layers of PPFR to assess its effectiveness for
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different concrete strengths. These specimens extend a previous study (Rousakis 2013) that
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targeted low concrete strength members, of around 15 MPa, confined by different FR
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reinforcements. Five specimens examined the dual confining effects of FRP jacket and FR
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wrapping. Two cylinders from each concrete batch were jacketed by 1 layer of GFRP and
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additionally wrapped by 3 layers of PPFR after resin curing. Another specimen in the C20 batch
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was GFRP jacketed and wrapped by 2 layers of PPFR. The remaining confined column of C20
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batch was wrapped by 3 layers of PPFR. These samples allowed for direct comparison of the
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FRP or the PPFR contribution to the hybrid confinement scheme.
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The technique does not use any mechanical means to stretch the rope during wrapping. The rope
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has a section area that can assure an adequate continuous hand tensile force on the rope during
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wrapping to avoid looseness of the rope. The ropes may have been reused several times in cases
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with no FR damage. Even in cases of extensive local fiber scratches, the healthy parts were
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connected with simple knots (see also Rousakis 2013). The use of the steel collars as mechanical
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anchorage for the ends of the continuous rope is ideal considering the small size of the
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specimens. Different pairs of collars with a diameter of 150-165 mm or 185-200 mm have been
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used. They had a width of 20 mm and a thickness of 1.5 mm. Both ends of the collar were
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connected with a simple mechanical device that used a threaded bolt (Figure 1b). Thus, the steel
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collars were open before application and they were adjusted to the desirable diameter. The rope
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ends were pushed on the concrete surface as the diameter of the top and bottom collars was
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reduced and the anchorage was accomplished through friction. Generally, the anchorage of the
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rope in real size columns will be easier to achieve. Anchorage of the rope outside the critical
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region of the member even without the use of suitable steel collars is possible. The beginning of
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the rope could be self-anchored by wrapping multiple layers outside the critical region. Then the
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rope could be wrapped around the column to provide the desirable confinement inside the critical
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region. Afterwards, the rope could continue outside the critical region and could be anchored
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through a simple knot on a fixed bolt. The external wrapping of concrete columns by PPFR
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needs no special treatment of the concrete surface or use of concrete surface coatings. Moreover,
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no resin impregnation of the FRs is required. Yet, large concave concrete substrates need suitable
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treatment since they may reduce significantly the confinement effectiveness. This restriction is
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valid for FRP or steel external wrapping as well. Therefore, the external FR wrapping constitutes
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a rather “directly applied and operating” strengthening technique when applied alone or after
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FRP jacketing.
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Test setup
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Four linear variable displacement transducers (LVDT) were used to measure the axial and lateral
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deformations of the cylinder. The first one measured the displacement between the loading
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platens and the second measured the distance between the steel collars. The lateral deformations
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of the concrete core were measured through two LVDTs, placed opposite to each other. A
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special laser displacement meter gave the relative deformation of the outer FR spiral between
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two target spots. Figures 1a, 2a and 2b present the test setup of a confined specimen, the attached
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LVDTs and the laser displacement meter.
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A loading machine of 3000 KN capacity was used for the compression tests. The discussions of
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the following sections suggest that the behavior of concrete with PPFR confinement may include
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a temporary drop of the bearing load after severe cracking of concrete (see also Rousakis 2013).
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A temporary load drop is evidenced after the fracture of GFRP, in cases of hybrid GFRP and
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PPFR strengthening. Therefore, the tests followed repeated axial compression – decompression –
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recompression cycles, manually, under increasing axial deformations. They included close cycles
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of around 0.05% compressive strain throughout the temporary load drop part and up to the point
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of load regaining. Then cycling was increased gradually to more than 1% strain. Two of the
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columns with 5 layers of PPFR confinement were subjected to monotonic axial compression.
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The strain rate of monotonic or cyclic loading was around 3*10-5 /sec. The tests of the specimens
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wrapped by PPFR (hybrid or not) were terminated before FR fracture. Table 2 presents the tested
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columns. The labels provide information regarding the specimens as follows. They denote the
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concrete batch (16 or 20), the confining material (G for GFRP and tPP for twisted polypropylene
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rope) and the number of layers (L1 for one layer). What follows is the mode of loading (that is M
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for monotonic or C for cyclic) and finally, the number of identical specimens.
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Experimental results
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Observed behavior
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Most of the tests that involved FR wrapping were stopped before reaching failure. The main
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reason for terminating the tests was the limitations of the loading machine as well as potential
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unsafe dislocations of the steel collars. However, in all those cases, the columns had already
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reached the critical point of load regaining after a temporary load drop. After that state, a severe
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disintegration of the concrete core occurred (referred to as concrete-related failure mode III and
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discussed in the following sections). The cylinders with GFRP jacketing presented an abrupt
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failure marked by the fracture of the sheet (denoted as failure mode I in what follows).
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Specimen 16tPPL5M1, wrapped by 5 layers of PPFR, displayed a “spring-like” behavior of the
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concrete at full decompression (Figure 2c). That mode of failure (mode II) was typical for fiber
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rope confinement (reported also in Rousakis 2013). The “spring-like” behavior of concrete is
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further explained in what follows. At high axial strains in the concrete, the core is extensively
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cracked. Still, it goes on supporting axial load because of the unbonded Fiber Rope. In each
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unloading event, the fiber rope is free to redistribute its restrictive action on the concrete core up
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to the point of balance with the remaining concrete lateral strains (plastic strains). After reaching
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the critical full unloading of the column, the weakened concrete core cannot attain a balance with
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the stressed FR. The FR squeezes the core inwards and gradually turns it into slices. This slicing
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of concrete core normal to the axis of the cylinder is related to the extensive multi-level cracking.
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For zero axial compression, the core is under transverse compression that varies among boundary
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confinement regions (additional confinement by the steel collars) and regions under rope
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confinement (at middle-height) as well as among regions with varying damage. The unbonded
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rope redistributes the lateral restriction freely and thus multiple concrete slices are formed. Then,
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the column presents an increase of its height, limited by the loading platens, described as an axial
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extension “spring-like” behavior. The specimen may come out higher than 300 mm (initial
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height) after full unloading. It can be clearly noted that in figure 2c the LVDT attached to the
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collars has lost contact with the metal base because of the lengthening of the specimen. The
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“spring-like” mode of failure reveals an optimized utilization of the strength and deformation
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capacity of concrete under rope confinement. Figure 2d shows the extensively cracked and sliced
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concrete core after the removal of the FR. Yet, the concrete core could bear further monotonic
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loading (or cyclic) if it was not fully unloaded. The full unloading was necessary because of the
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displacement limits of the compression machine.
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The tested specimens with full PP fiber rope wrapping of two, three or five layers presented no
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practical damage by the extensive cracking of concrete during compression, decompression and
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recompression even for multiple cycles. In cases of hybrid confinement, the PPFRs balanced the
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loss of lateral restriction after the first fracture of the GFRP jacket. PPFR wraps presented no
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damage after the testing and they were ready for reuse. The termination of some tests also
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involved the unsafe dislocation of steel collars, in high concrete axial strain (concrete-related
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mode of failure after FRP fracture, noted as mode I+III).
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Thorough observation of the GFRP sheet after FR removal revealed a different mode of fracture
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than in cases of common GFRP jacketing. The fracture of the specimen 16GL1R1 in Figure 3a
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was located in the middle height and in one side. It followed the typical mode of failure marked
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by the initiation of fracture at the point of higher strain concentration on the GFRP jacket, which
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instantaneously expanded towards the ends of the cylinders. This behavior of the FRP jacket is
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more or less common for different FRP materials, shape of concrete sections or mode of loading
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(Mirmiran et al 1998, Tamuzs et al. 2006, Rousakis and Karabinis 2008, Rousakis and Karabinis
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2012). On the other hand, after the removal of the PPFR, specimen 16GL1tPPL3R1 showed a
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widespread concrete cracking that resulted in the fracture of the GFRP sheet in multiple regions
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(3b and 3c). The FR redistributes the strains through friction with the GFRP sheet and the
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cracking of concrete leads to successive sheet fracturing. The GFRP after its initial fracture
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further restricts merely the concrete core through its interaction with the PPFR. Partial restrictive
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action by the FRP jacket after its first fracture is believed to be realized as the PPFR bridges the
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fractured FRP jacket through friction. At the same time, the PPFR further activates the rest of the
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jacket through friction as well. Successive loading and development of higher lateral strain,
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further activated the fractured GFRP jacket and multiple fractures occurred. Figures 3d and 3e
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present the same mode of GFRP sheet fracture for the specimen 20GL1tPPL2R1 after the
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removal of PPFR. Thus, in this proposed strengthening technique the capacity of the GFRP
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sheets can be further utilized since the damage of the jacket is more extensive. No “spring-like”
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concrete behavior happened in cylinders with hybrid confinement.
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The proper hybrid confinement design of the already applied GFRP jackets in columns with the
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addition of ultrahigh deformability elastic ropes may advance remarkably the deformational
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capacity of confined concrete after the FRP fracture. The FR controls the concrete expansion and
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thus concrete suffers an extensive spread of damage. The high degree of concrete damage is an
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indicator of the dissipated energy that could be vital in cases of seismic over loadings.
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Mechanical behavior
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This section focuses on the effect on concrete, from the use of PPFR as external confinement in
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already applied GFRP jackets. The performance of PPFR confined concrete is also presented for
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different concrete strengths and number of PPFR layers.
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The stress – strain curves for all tested columns are presented in Figures 4 to 12. Figure 4 shows
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the effect of the steel collars, mainly on the concrete strength of the cylinders, four and a half
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months after casting with respect to 28-day plain concrete strength. The concrete strength was
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higher as already discussed in materials’ section. Figures 5 and 8 illustrate the stress-strain
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behavior of glass FRP confined columns under cyclic loading. As expected, for higher concrete
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strength, the effectiveness of the same confinement was lower, in terms of normalized failure
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stress and strain.
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Figures 6, 9 and 10 present the specimens of different concrete strength confined by multiple
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layers of PPFR. C16 concrete confined by 5 layers of PPFR presented a negligible (5.7% load
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drop for 16tPPL5R1 and 1.8% for 16tPPL5R2) load drop with respect to the corresponding
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maximum stress at first stress reduction (labelled as fcc1 , Table 2). On the other hand, column
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16tPPL5M1 showed no temporary load drop (Figure 6). The monotonically loaded column
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presented a “spring-like” concrete failure at 6.8% axial strain and matching stress of 48.9 MPa.
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Similar was the behavior of the identical columns under cyclic loading reaching stresses between
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46 and 48 MPa at a strain around 6%. Those tests were terminated before failure to avoid
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undesirable collars’ dislocation. For higher concrete strength, the temporary load drop ranged
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between 19% and 23% (Figure 9). Even so, the load regaining part presented a higher rate of
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increase. All the specimens reached comparable bearing stresses in lower strain levels than
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similar columns with inferior concrete strength. The results of stress at 3% axial strain for the
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two groups of cylinders (see Table 2) show clearly this trend. Specimen C20tPPL3R1 displayed
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a temporary load drop of 43.2%, which is unacceptable for structural applications (Figure 10).
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Nevertheless, the FR managed to balance the temporarily unstable expansion of concrete and led
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the column to load regaining. Again, the rate of stress regaining was higher than in specimens
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with higher FR confinement. That behavior reveals another aspect of the extremely low
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sensitivity of the unbonded PPFR to stress concentrations or fragile concrete cracking through
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stress redistribution.
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Column C20tPPL3R1 presented the highest developed lateral strains for similar stress levels
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during loading.
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Hybrid glass FRP sheet – PPFR confinement of concrete
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Figures 7, 11 and 12 present the columns with hybrid GFRP and PPFR confinement. Proper
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design of the outermost PPFR wrapping can lead to a hybrid GFRP – PPFR strengthening of
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upgraded concrete performance under cyclic loading. All hybrid specimens presented an
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increased stress (fcc1 in Table 2) and strain at the initiation of GFRP fracture compared with the
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GFRP jacketed columns. The effect was higher for lower concrete strength. In fact, the initiation
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of FRP fracture at fcc1 was confirmed by the “plastic-like” sound heard after that characteristic
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point, marking the fracture of the FRP. Similar sounds were heard during successive cycles.
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Above evidence suggests that PPFR confinement does not hurt the GFRP jacket and does not
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cause premature failure.
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After the initiation of the GFRP fracture, the PPFR provided the dominant restrictive action to
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concrete. The abrupt GFRP jacket failures marked in Figures 5 and 8 were replaced by the
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smooth, controlled temporary softening behavior in Figures 7, 11 and 12. Proper design of the
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PPFR provided a temporary load drop for all hybrid specimens around 20%. The minimum load
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at softening was higher than the plain concrete strength for all the specimens. The effect was
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higher for lower concrete strength. Even specimen 20GL1tPPL2R1 with only 2 layers of PPFR
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displayed an acceptable load drop at softening. Yet, while the execution of multiple cycles
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resulted in a controlled behavior, the load regaining was significantly delayed compared with the
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column with 3 layers of PPFR. The stress - lateral strain behavior in Figures 11 and 12 suggests
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far better control of the lateral expansion of concrete by the heavier PPFR confinement. Figures
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7 and 11 reveal that for the heavy PPFR wrapping, the columns have regained the fcc1 stress level
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at around 4-5% axial strain. Consequently, PPFR confinement can provide a stable load
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recovery.
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Discussion
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Confinement effectiveness
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The comparative study of the behavior of PPFR confined columns and those with hybrid
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confinement reveals some interesting aspects of the techniques. The different confining
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techniques are assessed through the provided effective lateral rigidity (El). The El equals
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0.5kspρFR EFR for FRs where ρFR is the volumetric ratio of the fiber ropes (4t/d, t being the FR
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or FRP thickness and d the column section diameter). Parameter ksp is the effectiveness
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coefficient because of the clear spacing (s’ ) of the FR spiral, being ksp = 1-s’/2d (1 for full FR
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wrapping). The different values of the effective lateral rigidity for FRP and FR are presented in
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Table 2 (El,FR is the effective lateral rigidity for FRP and Elj for the FR). The GFRP jacketed
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columns 20GL1 and PPFR confined column 20tPPL3R1 in Figures 8 and 10 had almost the
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same effective lateral rigidity (El of 149.9 MPa and 176.3 MPa respectively). It seems that the
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GFRP jacket was activated earlier because of the use of impregnating resin, small thickness and
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high modulus of elasticity, compared with the PPFR wrapping. Most of the GFRP jacketed
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columns exhibited ever increasing bearing stress up to failure. On the other hand, the PPFR
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wrapping of similar lateral rigidity could reverse the softening concrete behavior only after
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0.65% strain and unacceptable load drop. Then, the stress regaining branch presented a very high
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rate of stress increase that reaches fcc1 stress at 2% strain. The dual FRP – FR confinement
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utilized the early FRP activation and the ultrahigh deformability of the PPFR. At the strain levels
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of the GFRP fracture, the FR has developed enough lateral pressure (restriction) to take over
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‘alone’. The FR could balance an abrupt FRP fracture with half of the initial lateral rigidity and
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provided a hardening response.
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The specimens with 5 layers of PPFR had similar El to the columns with hybrid confinement
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(around 300 MPa and 340 respectively) while it was twofold compared with the FRP jacketed
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columns. Again, the activation of the heavy PPFR confinement was more delayed. The bearing
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stresses up to the fracture of the GFRP were far higher in hybrid columns (or in columns
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jacketed exclusively by FRP) than in PPFR wraps. Nevertheless, the respective Figures show a
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different trend after the strain level of around 2.5%, after the GFRP fracture initiation and around
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the minimum bearing stress fcc2. At that strain level the bearing stresses of the PPFR confined
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columns were higher than in columns with hybrid confinement for C16 or C20 concrete batches
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(see also stress at 3% strain in Table 2).
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Proper design of hybrid confinement by FRP sheets and high deformability FR as the outermost
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reinforcement, utilized further the confining effects of the FRP sheet up to its fracture. Then,
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adequate FR additionally increased the strain ductility of concrete, which withstood also high
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axial loads and avoided an abrupt load capacity loss. Thus, any undesirable axial load capacity
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loss occurring in FRP or in steel stirrup confinement might be replaced by a smooth softening
350
behavior. Then, a hardening behavior could follow, of no “detected FR failure” for practical
351
applications in structural concrete members subjected to overload. The low sensitivity of the
352
tested ropes to local damage could be beneficial for highly energy dissipative seismic retrofitting
353
applications.
354
In cases of real columns with low axial load and moment force, the rope will be under tension in
355
the region where the column section is under compression. In the tension face of the column,
16
356
unbonded FRs may tend to loosen due to deformation compatibility issues and concrete tensile
357
crack formation. However, the proper design of the fiber rope dimensions and spacing as well as
358
the dilatation of the concrete core itself may counteract the aforementioned tendency. Certain
359
experimental results in the study by Shimomura and Phong (2007) show that unbonded rope
360
reinforcement could enhance efficiently the shear capacity and the displacement ductility of real
361
size columns under seismic load.
362
Modeling and design
363
Most available design-oriented confinement models depend the prediction of the ultimate stress
364
and strain values of concrete on the fracture of the confining means. Since the case of hybrid
365
GFRP and PPFR confinement does not include the fracture of the PPFR for the recorded tests,
366
this approach needs to be reconsidered. However, the existing models could predict the values at
367
the characteristic point of the fracture of the GFRP sheet for the case of dual hybrid confinement.
368
Two recent models are applied herein to predict the stress and strain at the distinctive point of
369
FRP fracture initiation (Teng et al. 2009, Rousakis et al. 2012, Rousakis et al. 2012b). The
370
empirical relations for the prediction of strength and strain of FRP confined columns by Teng et
371
al. (2009) acknowledge two decisive parameters. The first parameter is the confinement stiffness
372
and the second is the FRP tensile to concrete compressive strain ratio. The following relation
373
predicts the strength:
374
375
f cc 1 3.50 0.01 , when 0.01 2 j t j je f co , , , j 0.586 fu f co f cc co d 1, when 0.01 co f co
Relation (2) gives the strain at failure: 17
(1)
376
cc 0,80 1, 45 1,75 6,50 co ,
2 FRP t f co D
co
FRP co
(2)
,
377
In the research by Rousakis et al. (2012), the proposed strength model considers that the most
378
decisive parameter is the axial rigidity of the confining means (ρfEf ) normalized to the plain
379
concrete strength (fco) expressed as ρfEf /fco. According to the study, the effective lateral strain at
380
failure of concrete (εje) and the confinement effectiveness coefficient (k1) are considered variable,
381
in accordance with the experimental evidence. Yet, both of them were found strongly dependent
382
on the modulus of elasticity of the reinforcing fibers (Ef) of the confining means. Consequently,
383
the concrete strength prediction requires fewer parameters than in other proposed models. Only
384
the confinement volumetric ratio (ρf), the modulus of elasticity of the confinement reinforcing
385
fibers (Ef) and the plain concrete strength (fco) are necessary. The model has the following form:
386 387
fcc/fco = 1+ k1(fle/fco) = 1+ k1(0.5 ρf Ef εje/fco) = 1+ (ρf Ef/fco)( 0.5 k1 εje), that is fcc/fco = 1+ (ρfEf/fco)(α Ef 10-6 / Efμ + β)
(3)
388
with ρf = 4tf / d and Efμ = 10 MPa (for units’ compliance). For FRP sheet wraps α= -0.336 and
389
β= 0.0223. For FRP tube encased concrete α= -0.2300 and β= 0.0195.
390
The above models are applied to the experimental results of columns with dual GFRP-PPFR
391
confinement. The elaboration considers an equivalent thickness of GFRP jacket providing the
392
overall GFRP and PPFR confinement lateral rigidity. The analytical results are presented in
393
Table 3. The model by Teng et al. (2009) predicts the fcc1 and εcc1 values of concrete with an
394
absolute error of only 4.5% and 6.4% respectively. The strength model by Rousakis et al. (2012),
18
395
which is very sensitive to the modulus of elasticity of the reinforcing fibers, can also estimate the
396
fcc1 characteristic stress with an absolute error of only 6.5%.
397
The limited available results suggest that a normalized effective lateral rigidity of GFRP plus
398
PPFR, which is El,FR+j /f*co equal to 10, is adequate. In addition, the lateral rigidity of the PPFR
399
has to be equal or higher than that of the GFRP jacket. It results in an ever-increasing stress-
400
strain response of concrete, followed by an acceptable and controlled behavior of temporary load
401
drop (around 20%) upon initiation of GFRP fracture. Figure 13 presents the lower bound case of
402
20GL1PPL3R1. In terms of the minimum confinement stiffness ratio ρκ (as proposed by Teng et
403
al. 2009), a value of 0.02 for the hybrid scheme is adequate. Also the ρκ of the PPFR has to be
404
higher than 0.01. That way the crucial points of the complex stress-strain behavior of concrete
405
confined with the proposed hybrid scheme could be approached with the existing FRP models
406
for the lower bound adequate strengthening. Figure 13 also shows that the lower limit
407
strengthening secures a minimum load (at the softening part) higher than the plain concrete
408
strength. Moreover, the axial strain achieved for the maximum usable lateral strain of 0.4%, is
409
higher than the 1% axial strain (maximum usable strain according to ACI440-2008 and CNR-
410
DT200). Hence, the materials are fully utilized. The ultimate recorded stress is higher than the
411
fcc1 stress, denoting a full recovery. Finally, additional PPFR wrapping results in more than
412
double the strain ductility and the dissipated energy recorded up to the point of GFRP fracture
413
initiation (see the hatch-area).
414
Even though it is indirectly considered that the PPFR wrapping has similar effectiveness at early
415
loading stages with the FRP, - something that diverges from comparative experimental evidence
416
- the predictions are satisfactory. Most certainly, additional tests are necessary to provide reliable
417
design tools for this hybrid strengthening technique. 19
418
Conclusions
419
The paper presents the experimental investigation of a new hybrid FRP and PPFR confinement
420
applied on cylinders of different plain concrete strengths. The technique utilizes the unique
421
advantages of the PPFR confinement that needs no resin impregnation or mortar and the ends are
422
mechanically anchored. PPFR is resin-free and “green”, fully recycled material.
423
Adequate PPFR confinement of columns with normal concrete strength led to ever-increasing
424
bearing stresses or to a temporary controlled load drop. The specimens revealed ultimately a
425
“spring-like” concrete failure that was marked by the extensive disintegration of the concrete
426
core during full unloading. The corresponding axial strain levels at ultimate were as high as
427
6.8%. The behavior was similar for monotonic or cyclic loading.
428
As expected, for concrete of higher strength or for lower quantity of PPFR wrapping, the
429
effectiveness of the confinement in terms of ultimate concrete stress and strain was lower.
430
However, the load regaining part - after the temporary load drop - presented a higher rate of
431
increase. The FR managed to balance the temporarily unstable expansion of concrete and led to
432
load regaining even after an excessive load drop of 43.2%.
433
In cases of the hybrid confinement of concrete, a temporary load drop occurred upon initiation of
434
the fracture of the FRP sheet. Adequate FR withstood the energy that was released after the FRP
435
fracture. For higher imposed deformation, the system was stabilized and the load regaining
436
began. Thus, PPFR prevented an abrupt load capacity loss and ensured further increased strain
437
ductility of concrete and increased supported axial loads. The GFRP sheet contributed to the
438
lateral restriction of the concrete core even after its initial fracture. This was supported by the
20
439
fracture of the GFRP jacket in multiple regions. No concrete “spring-like” failure happened in
440
hybrid strengthening cases.
441
All the specimens with PP fiber rope wrapping, presented no fracture (or practical damage) even
442
after the extensive cracking and slicing of concrete or the multiple fracture of the GFRP and they
443
could be reused. The termination of the test in those cases was related to unsafe initiation of steel
444
collar dislocation in high axial concrete strain (except for the case of concrete-related “spring-
445
like” failure).
446
The proposed hybrid scheme may extend the upper limit applications of resin-involved FRP
447
jacketing (due to excessive required number of sheet layers or thickness) if the concrete strain
448
ductility is of importance.
449
Existing FRP models may reproduce the mechanical behavior of the hybrid strengthening, up to
450
the initiation of the FRP fracture. The analytical predictions of the Teng et al. (2009) model for
451
both fcc1 and εcc1 are accurate enough. A lower limit hybrid confinement stiffness of ρκ = 0.02 (with
452
PPFR having a ρκ ≥ 0.01) is required for an adequate overall response of the concrete columns.
453
However, more experiments are necessary on the field.
454
Acknowledgements
455
The author would like to acknowledge the contribution of Zarras S.A. for providing concrete, of
456
Sintecno S.A. for providing glass sheets and epoxy resins and Thrace Plastics Co. S.A. for
457
providing the polypropylene fiber ropes. Also, thanks are owed to undergraduate students Pekas
458
E., Dimitriadou T., Anezakis M. and Gouma M. for their help in the experimental program and
459
to the DUTh RC lab staff.
21
460
References
461
American Concrete Institute (2008). “Guide for the design and construction of externally bonded
462
FRP systems for strengthening concrete structures.” ACI 440.2R-08, Detroit.
463
Anggawidjaja D, Ueda T, Dai J, Nakai H. (2008). “Deformation capacity of RC piers by new
464
fiber-reinforced polymer with large fracture strain.” Cem Concr Compos. 2006;28:914–27.
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CNR-DT 200/2004, “Guide for the design and construction of externally bonded FRP systems
466
for strengthening existing structures,” National Research Council, Advisory committee on
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technical recommendations for construction, Rome July 13th 2004.
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Bournas, B.A.; Lontou, P.V.; Papanicolaou, C.G. and Triantafillou, T.C. (2007). “Textile-
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Reinforced Mortar (TRM) versus FRP Confinement in Reinforced Concrete Columns”, ACI
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Structural Journal, 104(6), pp. 740-748.
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Chen J.F., Li S.Q., Bisby L.A., Ai J. (2011). “FRP rupture strains in the split-disk test”, Composites: Part B 42, 962–972. Dai Jian-Guo, Bai Yu-Lei, Teng J. G. (2011). “Behavior and Modeling of Concrete Confined
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with FRP Composites of Large Deformability”, Journal of Composites for Construction, Vol.
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15, No. 6, pp. 963– 973.
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De Caso y Basalo F., Matta F., Nanni A. (2011). “Novel Test Method for Ultimate Hoop-Strain
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Characterization in FRP Jackets”, Journal of Materials in Civil Engineering, Vol. 23, No. 12,
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pp. 1633–1641.
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De Caso y Basalo F., Matta F., Nanni A. (2012). “Fiber reinforced cement-based composite
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system for concrete confinement”, Construction and Building Materials, 32 (2012) 55–65.
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De Lorenzis L, Tepfers R. (2003). “Comparative Study of Models on Confinement of Concrete
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Cylinders with Fiber-Reinforced Polymer Composites.” ASCE Journal of Composites for
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Construction, 7(3): 219-234.
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Fardis, M. N., Khalili, H. (1981). “Concrete encased in fibreglass-reinforced- plastic.” J. Am. Concr. Inst. Proc., 78(6): 440–446.
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Hu Y.M., Yu T., Teng J.G., (2011). “FRP-Confined Circular Concrete-Filled Thin Steel Tubes
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under Axial Compression”, Journal of Composites for Construction, Vol. 15, No. 5, pp.
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850–860.
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Karabinis A.I., Rousakis T.C., (2002). “Concrete confined by FRP material: a plasticity approach”, Engineering Structures Journal, Elsevier, 24:923–932. Liang M., Wu Z., Ueda T., Zheng J. and Akogbe R. (2012). “Experiment and modeling on axial
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behavior of carbon fiber reinforced polymer confined concrete cylinders with different sizes”,
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SAGE Journal of Reinforced Plastics and Composites, 31(6) 389–403.
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Matthys S.; Toutanji H.; Taerwe L., (2006). “Stress–Strain Behavior of Large-Scale Circular
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Columns Confined with FRP Composites”, Journal of Structural Engineering, Vol. 132, No.
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Mirmiran A., Shahawy M., Samaan M., El Echary H., Mastrapa J.C., Pico O. (1998). “Effect of
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Peled A., (2007). “Confinement of Damaged and Nondamaged Structural Concrete with FRP and
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TRC Sleeves”, Journal of Composites for Construction, ASCE, Vol. 11, No. 5,
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September/October 2007, pp. 514-522.
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Rousakis T.C., (2001). “Experimental investigation of concrete cylinders confined by carbon
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FRP sheets, under monotonic and cyclic axial compressive load”, Research Report, Chalmers
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University of Technology, Publ.01:2, Work No: 44. Göteborg, Sweden. 2001.
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Rousakis T.C., Karabinis A.I., Kiousis P.D., Tepfers R., (2008). “Analytical modelling of Plastic
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Behaviour of Uniformly FRP Confined Concrete Members”, Elsevier, Journal of Composites
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Part B: Engineering, Volume 39, Issues 7-8, October-December 2008, pp. 1104-1113.
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Rousakis T.C., Karabinis A.I., (2008). “Substandard Reinforced Concrete Members Subjected to
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Rousakis T.C., Karabinis A.I., (2012). “Adequately FRP confined reinforced concrete columns
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under axial compressive monotonic or cyclic loading”, RILEM Materials and Structures,
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Springer Netherlands 2012;45(7) 957-975. 23
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Rousakis T.C., Rakitzis T.D., Karabinis A.I. (2012). “Design - Oriented Strength Model for FRP
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Confined Concrete Members”, ASCE Composites for Construction 2012, Vol. 16, No. 6,
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pp.615-625.
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Rousakis T.C., Rakitzis T.D., Karabinis A.I. (2012b). “Empirical Modelling of Failure Strains of
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Uniformly FRP Confined Concrete Columns”, The 6th International Conference on FRP
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Composites in Civil Engineering – CICE 2012, 13 – 15 June, Rome.
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Rousakis, T. (2013). “Elastic Fiber Ropes of Ultra High Extension Capacity in Strengthening of
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Concrete Through Confinement”. J. Mater. Civ. Eng., accepted for publication (available at
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doi: 10.1061/(ASCE)MT.1943-5533.0000796).
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Scherer J (1999). “S&P—–Sintecno, FRP—polymer fibers in strengthening.” User guide, Brunnen. Shimomura Τ., Phong Ν. Η., (2007). “Structural Performance of Concrete Members Reinforced
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with Continuous Fiber Rope”, FRPRCS-8 Conference University of Patras, Patras, Greece,
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July 16-18, 2007.
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Shin M., Andrawes B. (2011). “Lateral Cyclic Behavior of Reinforced Concrete Columns
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Tamuzs V., Tepfers R., Chi-Sang You, Rousakis T., Repelis I., Skruls V., Vilks U., (2006).
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Yarns”, Mechanics of Composite Materials. Vol 42, No 1, ISSN 0191-5665. Springer
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Science+Business Media, Inc. pp. 13-32. Also in Russian. ISSN 0203-1272. pp. 21-44.
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Model for FRP-Confined Concrete”, ASCE Journal of Composites for Construction, 13(4),
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Triantafillou, T.C., Papanicolaou, C.G. (2005): “Textile Reinforced Mortars (TRM) versus Fiber
540
Reinforced Polymers (FRP) as Strengthening Materials of Concrete Structures”, SP230-6,
541
FRPRCS-7: Seventh International Symposium on FRP Reinforcement for Concrete
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Structures, ACI SP-230, Kansas City, USA, November 6-9 2005, pp. 99-118.
24
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25
Table Click here to download Table: RopesHybrid_TABLES_Rousakis20130215revised.doc
List of Table captions Table 1. Mechanical properties of glass FRP and of multi-strand fiber rope
Table 2. Experimental results
Table 3. Analytical predictions of Teng et al. 2009 and Rousakis et al. 2012 models
Page 1 of 4
Table 1. Mechanical properties of glass FRP and of multi-strand FR Label
Type of reinforcement
Area of Tensile cross Strength, section, fFRu AFR (mm2) (MPa)
Glass Fiber 0.154 Reinforced mm/m of Polymer width Z-twisted 2tPPFR strand 12.09 11gr/m polypropylene fiber rope
GFRP*
Tensile Modulus of Elasticity, EFR, (GPa)
Strain at failure εFRu, %
2044
73
2.8
405.3
2.0
20.4
*tensile properties of impregnated and cured sheet
Page 2 of 4
Table 2. Experimental results
Specimen label 16plain28avg 16plain_collarM1 16plain_collarM2 16plain_collar_avg 16GL1R1 16GL1R2 16tPPL5M1 16tPPL5R1 16tPPL5R2 16GL1tPPL3R1 16GL1tPPL3R2
Fiber Rope material
GFRP GFRP PP PP PP GFRP+PP GFRP+PP
Effective lateral Effective El,FR Effective FR FRP pressure lateral =fl'FR / confinement confinement fl'FR pressure εju ratio ke*ρFR ratio ρj (MPa) fl',j(MPa) (MPa)
0.00411 0.00411 0.307 0.296 0.304 0.190 0.190
0.00411 0.00411
4.20 4.20 62.2 60.0 61.5 38.4 38.4
4.20 4.20
20plain28avg 20plain_collarM1 20plain_collarM2 20plain_collar_avg 20GL1R1 GFRP 0.00411 4.20 20GL1R2 GFRP 0.00411 4.20 20tPPL5M1 PP 0.310 62.7 20tPPL5R1 PP 0.304 61.7 20tPPL5R2 PP 0.298 60.3 20tPPL3R1 PP 0.177 35.9 20GL1tPPL3R1 GFRP+PP 0.00411 4.20 0.194 39.3 20GL1tPPL3R2 GFRP+PP 0.00411 4.20 0.189 38.4 20GL1tPPL2R1 GFRP+PP 0.00411 4.20 0.132 62.7 *The results of the specimen are unreliable and are not included in the elaborations ** Failure mode I: FRP jacket fracture, failure mode II: concrete spring-like failure, ***In plain concrete specimens the strain values refer to maximum stress ****Bold values denote no Fiber Rope failure
305.5 294.9 302.3 188.8 188.7
Maximum Stress at first Stress Axial Strain Stress Drop Elj =fl',j / Reduction at fcc1, between fcc1 εju (MPa) fcc1(MPa) εcc1(%) and fcc2 % 20.5 0.180 22.4 0.443 27.7 0.358 25.1 0.401 35.8 149.9 1.305 149.9 48.9 0.0 6.817 27.6 5.7 0.401 27.7 1.8 0.397 40.4 16.1 149.9 1.635 39.3 24.6 149.9 1.724
149.9 149.9 308.2 303.0 296.2 176.3 193.0 188.7 130.9
149.9 149.9 149.9
25.8 34.6 32.8 33.7 38.7 36.1 32.4 34.9 30.5 33.0 42.6 42.9 42.7
0.269 0.352 0.426 0.389 1.129 0.784 0.387 0.375 0.401 0.407 1.633 1.420 1.445
failure mode III: concrete cracking
Page 3 of 3
0.8 11.9 23.0 20.8 19.0 43.2 21.0 18.5 17.8
Stress at 3% Ultim. Ultim. Ultim. strain Stress Axial Lateral fcc3 fcu Strain Strain (MPa) (MPa) εcu(%) εlu(%) 0.180 0.443 0.143 0.358 0.235 0.401 0.189 1.305 1.398 31.0 48.9 6.817 4.279 34.0 45.9 5.949 2.847 35.3 47.9 6.010 2.468 34.0 34.8 3.852 2.390 29.8 40.1 5.491 3.973
35.9 36.5 44.0 45.0 34.9 34.0
38.6 31.8 41.0 38.4 47.1 35.4 45.2 46.3 36.8
0.269 0.352 0.426 0.389 1.498 1.238 3.890 3.460 3.333 2.140 4.825 4.979 4.034
0.196 0.381 0.288 1.350 1.498 1.117 3.081 2.820 2.735 2.100 4.243
failure mode**
I I* II III (early stop) III (early stop) I + III (early stop) I + III (early stop)
I I III (early stop) III (early stop) III (early stop) III (early stop) I + III (early stop) I + III (early stop) I + III (early stop)
Table 3. Analytical predictions of Teng et al. 2009 and Rousakis et al. 2012 models Teng et al. 2009 Absolute error of strength εcc1 anal / predictions εcc1 exper
Absolute error of strain predictions
Rousakis et al. 2012 Absolute fcc1 anal error of / strength fcc1 exper predictions
Specimen label
El,FR
tj equiv.
fcc1 anal / fcc1 exper
16GL1tPPL3R1 16GL1tPPL3R2
338.7 338.6
0.348 0.348
0.92 0.95
0.077 0.050
1.11 1.06
0.113 0.056
0.95 0.98
0.047 0.020
20GL1tPPL3R1 20GL1tPPL3R2 20GL1tPPL2R1
342.9 338.6 280.8
0.352 0.348 0.288
1.02 1.01 0.94
0.025 0.013 0.060
0.98 1.12 1.01
0.025 0.116 0.010
1.11 1.10 1.05
0.109 0.098 0.050
Average
0.97
0.045
1.05
0.064
1.04
0.065
Page 3 of 3
Figure1 Click here to download Figure: RopesHybrid_790rev_FIGURE1.pdf
Collar Lateral 1
Axial 2 Laser target spots
Rope
Axial 1 (a)
1
(b)
Lateral 2
(c)
Figure2 Click here to download Figure: RopesHybrid_790rev_FIGURE2.pdf
(a)
2
(b)
(c)
(d)
Figure3 Click here to download Figure: RopesHybrid_790rev_FIGURE3.pdf
(a)
3
(b)
(c)
(d)
(e)
50
50 45 40 35 30 25 20 15 10 5 0
45
20plain_collar
30 25 20 15 10 5
0
Lateral Strain, εl
(b)
0 0. 00 2 0. 00 4 0. 00 6 0. 00 8
0
-0 8 .0 2 -0 .0 18 -0 .0 16 -0 .0 14 -0 .0 12 -0 .0 1 -0 .0 08 -0 .0 06 -0 .0 04 -0 .0 02
06
16plain_collar
0. 0
04
0. 0
0. 0
0. 0
02
16plain28
Axial Strain, εa (a)
4
0.4% lateral strain
40 35
20plain28
0
Axial Stress, fa (MPa)
Figure4 Click here to download Figure: RopesHybrid_FIGURES_edit4.pdf
Axial Strain, εa
Figure5 Click here to download Figure: RopesHybrid_FIGURES_edit5.pdf
50 45
Axial Stress, fa (MPa)
40
1% axial strain 0.4% lateral strain 16GL1R1
35 30 25 20 15 10 5 0 -0.05 -0.04 -0.03 -0.02 -0.01
Lateral Strain, εl
5
0
0.01
0.02
0.03
0.04
Axial Strain, εa
0.05
0.06
0.07
Figure6 Click here to download Figure: RopesHybrid_FIGURES_edit6.pdf
50
Axial Stress, fa (MPa)
1% axial strain
0.4% lateral strain
45
fcc1
40
fcu
fcc2
35 30 25 20 15
16tPPL5R1 T9-D
10
Τ-16 16tPPL5M1
5
Τ-15 16tPPL5R2
0
-0.05 -0.04 -0.03 -0.02 -0.01
Lateral Strain, εl
6
0
0.01
0.02
0.03
0.04
Axial Strain, εa
0.05
0.06
0.07
Figure7 Click here to download Figure: RopesHybrid_FIGURES_edit7.pdf
50 45
0.4% lateral strain
1% axial strain fcc1
Axial Stress, fa (MPa)
40
fcu
fcc2
35 30 25 20 15 10
16GL1tPPL3R1 ∆8_16GR/PP3 ∆9_16GR/PP3 16GL1tPPL3R2
5 0 -0.05 -0.04 -0.03 -0.02 -0.01
Lateral Strain, εl
7
0
0.01
0.02
0.03
Axial Strain, εa
0.04
0.05
0.06
0.07
Figure8 Click here to download Figure: RopesHybrid_FIGURES_edit8.pdf
50
1% axial strain
0.4% lateral strain
45 40
Axial Stress, fa (MPa)
35 30 25 20
20GL1R1 ∆3_20GR ∆4_20GR 20GL1R2
15 10 5 0 -0.05 -0.04 -0.03 -0.02 -0.01
Lateral Strain, εl
8
0
0.01
0.02
0.03
0.04
Axial Strain, εa
0.05
0.06
0.07
Figure9 Click here to download Figure: RopesHybrid_FIGURES_edit9.pdf
50
0.4% lateral strain
1% axial strain
45 40
Axial Stress, fa (MPa)
35 30 25 20 15
20tPPL5M1 PP2tj_1
10
20tPPL5R1 T2-BC+AB/2
5
T1220tPPL5R2
0 -0.05 -0.04 -0.03 -0.02 -0.01
Lateral Strain, εl
9
0
0.01
0.02
0.03
0.04
Axial Strain, εa
0.05
0.06
0.07
Figure10 Click here to download Figure: RopesHybrid_FIGURES_edit10.pdf
50 45 40
1% axial strain 0.4% lateral strain 20tPPL3R1
Axial Stress, fa (MPa)
35 30 25 20 15 10 5 0 -0.05 -0.04 -0.03 -0.02 -0.01
Lateral Strain, εl
10
0
0.01
0.02
0.03
0.04
Axial Strain, εa
0.05
0.06
0.07
Figure11 Click here to download Figure: RopesHybrid_FIGURES_edit11.pdf
50
0.4% lateral strain
1% axial strain
45
Axial Stress, fa (MPa)
40 35 30 25 20 15 10
20GL1tPPL3R1 ∆9_20GR/PP3 20GL1tPPL3R2 ∆10_20GR/PP3
5 0 -0.05 -0.04 -0.03 -0.02 -0.01
Lateral Strain, εl
11
0
0.01
0.02
0.03
0.04
Axial Strain, εa
0.05
0.06
0.07
Figure12 Click here to download Figure: RopesHybrid_FIGURES_edit12.pdf
50 45
1% axial strain
0.4% lateral strain
40
20GL1tPPL2R1
Axial Stress, fa (MPa)
35 30 25 20 15 10 5 0 -0.05 -0.04 -0.03 -0.02 -0.01
Lateral Strain, εl
12
0
0.01
0.02
0.03
0.04
Axial Strain, εa
0.05
0.06
0.07
Figure13 Click here to download Figure: RopesHybrid_FIGURES_edit13.pdf
50 45
0.4% lateral strain
1% axial strain
Axial strain at 0.4% lateral strain
fcc1 stress level
40
< 20% fcc1 f*co stress level
30 25 20 15 10 5 0 -0.03
20GL1tPPL3R1 -0.02
-0.01
Lateral Strain, εl
13
dissipated energy attributed to PPFR wrap
Axial Stress, fa (MPa)
35
0
0.01
0.02
0.03
Axial Strain, εa
0.04
0.05
Figure Captions List
List of figure captions Fig. 1. Layout of FR confinement and position of the steel collars (a). Steel collar (b). Z-twisted two-strand polypropylene fiber rope (c).
Fig. 2. Characteristic photos of specimen 16tPPL5M1. Test setup and instrumentation with LVDTs (a) and laser meter (b). ‘Spring-like’ behaviour at full decompression (c). Extensively cracked and sliced concrete core after rope removal (d).
Fig. 3. Specimen 16GL1R1 after failure. One-sided GFRP fracture (a). Specimen 16GL1tPPL3R1 after failure. Multiple fracture regions (b) and (c). Two – sided GFRP fracture of specimen 20GL1tPPL2R1 (d) and (e).
Fig. 4. Axial stress versus axial and lateral strain curves of plain concrete cylinders and columns with steel collars under monotonic axial load
Fig. 5. Axial stress versus axial and lateral strain curves of C16 concrete cylinder wrapped by 1 layer of glass FRP under cyclic axial load
Fig. 6. Axial stress versus axial and lateral strain curves of C16 concrete cylinders wrapped by 5 full layers of polypropylene fiber rope, under monotonic or cyclic axial load
Fig. 7. Axial stress versus axial and lateral strain curves of C16 concrete cylinders wrapped by 1 layer of glass FRP and 3 full layers of polypropylene fiber rope, under cyclic axial load
Fig. 8. Axial stress versus axial and lateral strain curves of C20 concrete cylinders wrapped by 1 layer of glass FRP under cyclic axial load
Fig. 9. Axial stress versus axial and lateral strain curves of C20 concrete cylinders wrapped by 5 full layers of polypropylene fiber rope under monotonic or cyclic axial load
Fig. 10. Axial stress versus axial and lateral strain curves of C20 concrete cylinder wrapped by 3 full layers of polypropylene fiber rope under cyclic axial load
Fig. 11. Axial stress versus axial and lateral strain curves of C20 concrete cylinders wrapped by 1 layer of glass FRP sheet and 3 full layers of polypropylene fiber rope, under cyclic axial load
Fig. 12. Axial stress versus axial and lateral strain curves of C20 concrete cylinder wrapped by 1 layer of glass FRP and 2 full layers of polypropylene fiber rope, under cyclic axial load
Fig. 13. Characteristic design milestones and performance of the lower bound adequate stressstrain behaviour for columns with hybrid GFRP-PPFR confinement (20GL1tPPL3R1 case).
