The 7 th International Conference on FRP Composites in Civil Engineering International Institute for FRP in Construction
FIBER-REINFORCED POLYMER STRENGTHENED STEEL REINFORCED MASONRY WALLETTES IN OUT-OF-PLANE BENDING Bowen SHEN Graduate Research Assistant, North Carolina State University, USA
[email protected]
Rudolf SERACINO Associate Professor, North Carolina State University, USA
[email protected]
ABSTRACT: Research on the strengthening of masonry structures subject to out-of-plane bending has focused exclusively on unreinforced masonry comprised of concrete masonry units (CMU), clay brick, or natural stone. Consequently, major international design guides are limited in their scope to unreinforced masonry. However, a significant portion of today’s masonry building stock consists of steel reinforced CMU walls including, for example, residential construction in coastal areas prone to extreme hydrometeorological events and earth retaining walls, both of which are subject to out-of-plane loading. The purpose of the present research is to create the first database of steel reinforced CMU wallettes strengthened with externally bonded (EB) fiber-reinforced polymer (FRP) systems subject to out-of-plane bending. The experimental program comprises nine steel reinforced CMU wallettes, one of which served as an unstrengthened control, while the other eight were strengthened with wet lay-up Glass or Carbon EB-FRP systems. All the wallettes were simply-supported and tested in three-point out-of-plane bending. Two failure modes were observed: FRP debonding; and FRP rupture. The experimental results show that the load carrying capacity and post-cracking stiffness of all the strengthened wallettes increased. This paper summarizes the results of the experimental program and compares the experimentally measured maximum FRP strains with the prediction of the same using the recommendations of ACI 440.7R-10.
1. Introduction Steel reinforced concrete masonry unit (CMU) walls are a typical type of wall system, which is widely used in residential construction in coastal regions prone to extreme hydro-meteorological events, and earth retaining structures. Strengthening may be required due to a change in occupancy or increases in allowable loading. Similarly, repair may be required due to environmental deterioration or damage. It is well-known that externally bonded (EB) fiber-reinforced polymer (FRP) systems are an efficient means of improving the out-of-plane flexural behavior of masonry structures. However, previous research and current design guides (for example, Albert and Elwi (2001), Tumialan and Galati (2003), and ACI 440.7R (2010)) focus exclusively on the FRP strengthening of unreinforced masonry (URM) walls. Hence, the purpose of this research is to begin to address this gap by creating an experimental database of EB-FRP strengthened steel reinforced CMU wallettes subject to out-of-plane bending. The observed failure modes and flexural behavior of the strengthened walls may then be compared to the predictions of existing design recommendations originally intended for URM. The experimental program consisted of nine steel reinforced CMU wallettes, including one unstrengthened control wallette. The other eight wallettes were strengthened with wet lay-up Glass or Carbon EB-FRP systems designed to fail by the intermediate crack (IC) debonding mechanism so that existing FRP debonding strain models, such as that provided in ACI 440.7R (2010), can be validated by the experimental data.
2. Experimental Program 2.1. Wall Construction Nine wallettes were constructed using lightweight two-cell CMUs. The nominal width and height of the wallettes was 620 mm (24.5 in.) and 2030 mm (80 in.), respectively, with a single wythe thickness of 200 mm (8 in.). The wallettes were built by professional brick layers in a running bond pattern using one and a half CMUs per course. The center core of the cross-section was fully grouted with concrete and reinforced using a single 16 mm (No. 5) steel reinforcing bar over the full height of the wallette. To control shrinkage, 3.8 mm (9 gage) ladder-type wire joint reinforcement was used with a vertical spacing of 400 mm (16 in.).
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As is typically done in practice, a 13 mm (No. 4) steel reinforcing bar was located horizontally in the top course. In addition, the outside cores of the top and bottom two courses were filled with concrete to avoid a premature shear sliding failure near the supports. 2.2. FRP Installation The wallette surface was prepared according to the manufacturer’s instructions to ensure a good quality bond. Sharp edges and projections on the wallette surface were removed using a grinding disc, and the dust generated was removed using high-pressure air. In a typical wet lay-up procedure, the pre-cut fabric was saturated with epoxy on a clean horizontal surface. In the meantime, another batch of epoxy was thickened using Cab-o-sil and applied to the wallette to prime the surface and fill voids. The saturated fabric was then applied to the prepared wallette surface in the vertical orientation taking care to ensure that the fibers were taut and properly aligned. For the wallettes strengthened with two layers of FRP, the second saturated fabric layer was immediately applied directly to the first layer in a similar manner. The strengthened wallettes were then fully covered using plastic sheeting to protect the FRP from the environment during curing. All strengthened wallettes were allowed to cure for at least 10 days prior to testing. The typical material properties of the Glass FRP (GFRP) and Carbon FRP (CFRP) strengthening systems, as provided by the manufacturer, are summarized in Table 1. Table 1 – Manufacturer’s properties of fibers and laminates. Thickness Tensile strength Elastic modulus Rupture strain Material (MPa) (GPa) (mm) () Carbon fiber 0.33 4480 234.4 19000 CFRP laminate 1.27 966 69.6 12000 Glass fiber 0.345 2275 72.4 40000 GFRP laminate 1.1 552 27.6 22000 2.3. Test Setup All of the wallettes were simply supported over a vertical span of 1670 mm (66 in.). Two steel tubes were tied back to a reaction wall by steel tension bars to provide the boundary supports. The FRP was 1525 mm (60 in.) long so that the ends were not clamped by the supports. A hydraulic jack was located horizontally at the mid-height of the wallettes to apply the concentrated load, which was increased monotonically to failure. A steel tube was located between the hydraulic jack and the wallette to distribute the load over the full width. A sheet of 6.35 mm (0.25 in.) thick rubber was inserted at all of the interfaces of the steel tubes and the wallette to evenly distribute the load along the full width of the wallette and minimize any stress concentrations due to unevenness. A steel roller was used at the bottom of the wallette so that it may rotate freely during testing. Figure 1 shows an overview of the test setup. 2.4. Instrumentation A load cell was attached to the hydraulic jack to record load continuously during the test. Two string pots were located at mid-height at both vertical edges of the wallette to measure the maximum out-of-plane displacement. A string pot was also located at the top and bottom supports to monitor the displacement of the supports so that the rigid body movement of the wallette may be subtracted from the measured midheight displacement. A 6.35 mm (0.25 in.) long linear strain gage was bonded vertically on the surface of FRP at the mid-height bed joint to measure the expected maximum FRP tension strain. In addition, an advanced non-contact 3D position measurement system (Optotrak) was used to validate and supplement the data recorded from the traditional strain and displacement instrumentation. Figure 2 shows the setup of the Optotrak system prior to a test. The camera records the 3D position data from each LED target that can be used to calculate continuous out-of-plane displacement and FRP strain distributions along the full height of the wallettes. As can be seen in Fig. 2, the LED targets were located uniformly along the height of the FRP composite with a center-to-center spacing (gage length) of 50 mm (2 in.).
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Figure 1 – Overview of Test Setup
Figure 2 – Overview of Optotrak System Setup
3. Test Results Table 2 summarizes the primary experimental results of each wallette. For reference, the strengthened wallettes are designated with two letters followed by a number. The first letter represents the type of FRP: “C” for CFRP; and “G” for GFRP. The second letter represents the number of FRP layers: “S” for single layer; and “D” for double layer. The number represents the width of the FRP laminate in mm. Thus, CS150 represents a wallette strengthened with one layer of 150 mm (6 in.) wide CFRP. It should be noted that the wallette strengthened with one layer of 75 mm (3 in.) wide GFRP, GS75, failed by rupture, which was not initially expected. Hence, this test was repeated and these two wallettes are designated as GS75-1 and GS75-2. It should also be noted that the FRP strain values given in Table 2 are the maximum recorded strain calculated using the Optotrak data, which is the average strain over the 50 mm (2 in.) gage length, and the deflection is that corresponding to the ultimate load. Also shown for reference in Table 2 is the FRP reinforcement ratio for each strengthened wallette. These were selected to be within the representative range of values used to derive the design procedure in ACI 440.7R (2010). Two failure modes were observed for the strengthened wallettes: (1) IC debonding of the FRP; and (2) FRP rupture. Figure 3 shows representative photographs of these failure modes.
Figure 3 – Typical Failure Modes
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Specimen Control CS75 CS150 CS200 CD75 GS75-1 GS75-2 GS150 GD75
FRP reinforcement ratio (%) N.A. 0.077 0.154 0.204 0.154 0.066 0.066 0.133 0.133
Table 2 – Experimental results Maximum FRP Ultimate load strain (kN) () 20.0 N.A. 39.2 8313 48.8 6751 57.8 5984 44.8 7751 26.8 13479 23.2 14172 28.7 11782 34.9 14936
Mid-Height Deflection (mm) 71.5 19.0 16.3 16.3 12.8 9.1 14.0 12.2 24.1
Failure mode Flexure IC Debonding IC Debonding IC Debonding IC Debonding FRP Rupture FRP Rupture FRP Rupture IC Debonding
4. Brief Discussion of Results The control wall failed in a ductile tension controlled flexural mode and the test was stopped prior to complete collapse. All of the CFRP strengthened wallettes and GD75 failed by IC debonding, and the overall behavior was as expected. Interestingly, all of the single layer GFRP strengthened wallettes failed by FRP rupture. Figure 4 shows the load – mid-height deflection curves for all of the wallettes, including the control wall. 60
CS200 CS150
50 Load (lkN)
CD75
40
CS75 GD75
30
GS150 GS75-1
20
GS75-2 Control
10 0 0
10
20
30
40
50
60
70
Out-of-plane mid-height deflection (mm) Figure 4 – Load – Mid-Height Deflection Curves The experimental results show that the out-of-plane flexural capacity and post-cracking stiffness increased for all of the strengthening alternatives considered, with that of the CFRP strengthened wallettes being higher than the GFRP strengthened wallettes. And as may be expected, the load carrying capacity and post-cracking stiffness are related to the reinforcement ratio of each FRP strengthening alternative. It is interesting to observe that while the maximum recorded FRP strain is slightly less for GS75-1 compared to that of GS75-2, the load carrying capacity of GS75-1 is higher. After testing, some of the masonry and grouted concrete were removed from the wallettes to observe the condition of the crosssection at the maximum moment region and to measure the actual location of the internal steel reinforcing bar. While the steel reinforcement was intended to be centrally located in the cross-section, it was found to be closer to the tension face in GS75-1 and closer to the compression face in GS75-2, which may
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account for the variation in flexural capacity observed. This is confirmed in Fig. 4, where it can be seen that the flexural strength of GS75-1 after FRP rupture is higher that of GS75-2 and the control wallette. Further, it is noted that the measured maximum FRP strain for the GS wallettes were at best only 75% of the expected rupture strain, even though all these wallettes failed by FRP rupture. It is speculated that the sharp edge at the masonry flexural crack face may be locally exacerbating the stress concentration resulting in a perceived premature rupture failure. For illustrative purposes, Fig. 5 shows the FRP strain distribution for GD75 at selected applied loads up to failure by IC debonding. The increase in FRP strain below mid-height as the ultimate load was reached is an indication of the rapid propagation of the IC debonding crack from the location of maximum moment towards the bottom of the wallette. It is also interesting to note that the local FRP strain maxima tend to occur at the location of the bed joints where the masonry flexural cracks develop. 1.8
vertical location (m)
1.6 1.4 1.2 1 10 kN
0.8
20 kN
30 kN
35 kN
0.6 0.4 0.2 0
5000
10000
15000
FRP strain () Figure 5 – FRP Strain Distribution for Wallette GD75
5. Comparison with ACI 440.7R-10 According to ACI 440.7R (2010) the effective (maximum) FRP strain, fe, is given by
fe m *fu
(1)
where m is a bond-dependent coefficient for flexure which is equal to 0.45 for any EB-FRP system, and fu* is the ultimate rupture strain of the FRP as reported by the manufacturer. From Eq. 1, fe for the CFRP used in this test program is 5400 , and the same for the GFRP is 9000 , both of which are less than all of the maximum FRP strains recorded (see Table 2) for the respective strengthening material. It must be noted, however, that for all the strengthening systems considered in this test program the recommended limit on the allowable force transferred per unit width, pfm, given by the following, is exceeded.
p fm nt f f fe 260 N mm
(2)
where n is the number of plies of FRP laminates, tf is the nominal thickness of one ply of the FRP reinforcement, and ffe = Effe is the effective stress level in the FRP reinforcement where Ef is the tensile modulus of the FRP.
6. Conclusions A test series comprising nine steel reinforced CMU wallettes loaded in three-point out-of-plane bending was undertaken, eight of which were strengthened with either Glass or Carbon externally bonded FRP systems using the wet lay-up technique. The observed failure modes were IC debonding and FRP rupture. All the strengthened wallettes exhibited an increase in load carrying capacity and post-cracking stiffness
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compared to that of the unstrengthened control wallette, with a corresponding reduction in out-of-plane displacement regardless of the failure mode. It is believed that this represents the first experimental test program of FRP-strengthened steel reinforced CMU walls in out-of-plane bending. Further, it is shown that the measured maximum FRP strain in all cases exceeded the prediction provided by ACI440.7R (2010), even though the bond-dependent coefficient for flexure was derived from tests on URM. Even though the outcomes suggest the current design recommendations may be safely used, it should be noted that the force per unit width that the FRP system transfers to the masonry substrate exceeded the recommended limit provided in all cases.
7. Acknowledgements The authors would like to thank Pinnacle Masonry, Edge Structural Composites, and the Concrete Reinforcing Steel Institute for donating the materials used in this experimental program. The technical input from Ravi Kanitkar, Gustavo Tumialan, Danks Burton and Travis Cole, as well as the technical assistance provided by the staff of the Constructed Facilities Laboratory at North Carolina State University, are all greatly appreciated.
8. References ACI440.7R “Guide for the Design and Construction of Externally Bonded Fiber-Reinforced Polymer Systems for Strengthening Unreinforced Masonry Structures”, American Concrete Institute, Farmington Hills, MI, 2010, 50 pp. ALBERT, Michael, ELWI, Alaa, “Strengthening of Unreinforced Masonry Walls Using FRPs”, ASCE Journal of Composites for Construction, Vol. 5, Issue 2, May 2001, pp. 76-84. TUMIALAN, Gustavo, GALATI, Nestore, “Fiber-Reinforced Polymer Strengthening of Unreinforced Masonry Walls Subjected to Out-of-Plane Loads”, ACI Structural Journal, Vol. 100, No. 3, MayJune 2003, pp. 321-329.
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