The most efficient and popular application of fiber-reinforced polymer (FRP) materials is to use them to confine reinforced concrete (RC) columns for strength and ductility enhancement. To achieve a reliable design for FRP strengthening of such columns under the action of axial loading or combined axial and lateral loading, an in-depth understanding of the compressive stress-strain behavior of FRP-confined concrete is very important. At present, a large number of experimental studies have been conducted particularly on FRP-confined circular concrete columns. Theoretical stress-strain models, either design-orientated ones (e. g. Lam and Teng 2003) or analysis-orientated ones (e. g. Jiang and Teng 2007), have been developed to predict the compressive strength, the ultimate compressive strain, as well as the full-range compressive stress-strain behavior of FRP-confined concrete. However, no experimental work has been conducted on the compressive stress-strain behavior of concrete confined with LRS FRPs. It remains unclear if the existing confinement models are applicable for LRS FRP-confined concrete.
Given the above context, compressive tests were conducted on a total of 42 cylindrical concrete specimens, comprising 36 FRP-wrapped specimens and 6 control specimens, which were prepared and tested under monotonic uni-axial compression (Dai et al. 2011). Each specimen was 152 mm in diameter and 305 mm in height. Among the 36 specimens, 9 specimens were wrapped with AFRP FRP jackets, other 9 specimens were wrapped with PET FRP jackets, and the rest were wrapped with PET FRP jackets. AFRP jackets were used in the test program because the existing compressive tests on AFRP-confined concrete are very limited and the rupture strain
Fig. 14.2 Typical failure of FRP-confined concrete (a) AFRP (b) PEN FRP (c) PET FRP
of AFRP lies between those of CFRP and PEN FRP (see Fig. 14.1), representing a transition state of rupture strain from high modulus FRPs to low modulus FRPs.
It was found that the typical failure of LRS FRP-confined concrete was still due to the hoop tensile rupture of the FRP jacket outside the overlapping zone as observed in conventional FRP-confined concrete (see Fig. 14.2). But the failure at the ultimate state was smooth and quiet, unlike the explosive one observed in AFRP-confined concrete. The average hoop rupture strain of FRP jackets from these compression tests was found to be 3.0%, 4.5% and 7.5%, respectively, for AFRP, PEN FRP and PET FRP, while most existing tests on concrete confined by GFRP and CFRP exhibited hoop rupture strains of around 2% and 1%, respectively (Lam and Teng 2004). Similar to those of conventional FRP-confined concrete with a sufficient level of FRP confinement, the stress-strain curves of LRS FRP-confined concrete also exhibited a monotonically ascending bilinear shape with rapid softening in a transition zone around the stress level of unconfined concrete strength. Both the compressive strength and the ultimate axial strain were significantly enhanced (see Fig. 14.3).
To see if existing confinement models, which have been developed based on tests of conventional FRP-confined concrete, can describe appropriately the compressive behavior of LRS FRP-confined concrete, an analysis-orientated model, which was proposed by Jiang and Teng (2007) and was reported to be the most accurate among similar models, was used to predict the experimental compressive stress-strain curves of LRS FRP-confined concrete following an incremental iteration calculation process (Dai et al. 2011). This analysis-oriented model is, in principle, applicable to concrete confined by all sorts of materials as long as the relationship between the hoop strain (or lateral strain) and the confining pressure of the confining jacket is established. The predictions from Jiang and Teng’s model are in close agreement with the test results in terms of the compressive strength, but overestimate the ultimate axial strain of LRS FRP-confined concrete as shown in
Fig. 14.3 Compressive stress-strain relationships of LRS FRP-confined concrete
Fig. 14.3, in which PET FRP-confined concrete has been taken as an example. This discrepancy indicates that the lateral strain-axial relationship in Jiang and Teng’s model is inaccurate for strains larger than that normally experienced by conventional FRP jackets. Based on regression of the test results, a new lateral strain — axial strain equation was proposed for LRS FRP-confined concrete as follows (Dai et al. 2011):
^ = (1 + 8 — )-[1.024(—)0350 + 0.089(—)] (14.1)
Єсо fco Є co Є co
where e is the axial strain of concrete; e is the lateral strain of concrete; e is the axial strain at the peak axial stress of unconfined concrete; sl is the lateral confining pressure; and fco’ is the compressive strength of unconfined concrete.
Based upon the proposed model, a simple comparison among concrete confined with FRP jackets made of five different fiber sheets (CFRP, GFRP, AFRP, PEN and PET) was also made as shown in Fig. 14.4 (Dai et al. 2011). A concrete cylinder with an unconfined concrete strength of 38 MPa was assumed for such a comparison and the rupture strain of fibers used in the analysis were assumed to be 0.095%, 1.7%, 3.0%, 4.5%, and 7.5%, which are typical values for CFRP, GFRP, AFRP, PEN FRP, and PET FRP, respectively. Provided that all FRP-confined concrete achieves the same energy absorption, which is evaluated using the area underneath the compressive stress-strain curve, at the rupture of the fiber sheet, the needed tension stiffness of PEN FRP and PET FRP jackets are found to be 19.0% and 13.3%, respectively, of that of CFRP jackets. Apparently, LRS FRP jackets lead to a more economic and attractive confinement solution in terms of the energy absorption capacity.
Fig. 14.4 Energy absorption comparison of conventional and LRS FRP-confined concrete