EXPERIMENTAL RESULTS

Transcription

1 Chapter 4 EXPERIMENTAL RESULTS 4.1 Introduction This chapter presents the results from the half scale interior Corcon rib beam-column subassemblage and the FRP repaired subassemblage. As described in chapter 3, the original undamaged specimen was detailed according to the current Australian design practice with no special provision for seismicity. The repaired specimen was detailed to cover detailing problems identified in the original specimen. The test results and various observations related to the test are presented in this chapter. The overall performance of the specimen is assessed and discussed in terms of strength, stiffness, energy dissipation, ductility and displacement capacity st interior specimen Observed behaviour General There were no visible cracks in the specimen before the test and after applying the initial gravity loading. All new cracks and extensions of old cracks were numbered according to the cycle number in which they were first seen. The specimen appeared to perform fairly well under lateral deformations up to 3% drift level. The cracks found on the specimen 88

2 were mainly cracks due to flexure on the beam and column. There were no cracks appeared due to secondary effects Types and formation of cracks As mentioned previously all cracks were numbered according to the cycle number in which they were first seen. However, there is a possibility that some cracks may have formed in an earlier cycle but were too fine to be detected. Cracks formed under positive loading were marked in red, while those formed under negative loading were marked in blue. The direction of loading direction (positive and negative) was indicated in figure 3-2 of chapter 3. Figure 4-1 illustrates the cracks found in the specimen after completing all loading cycles. This gives a clear picture of the overall cracking pattern of the specimen. 89

3 C4-.1 C6-.1 S C4-.2 C5-.1 C5-.1 C4-.2 C4-.1 C8-.2 C4-.2 C4-.1 C4-.1 C6-.1 N C4-.3 C7-.6 East Elevation C4-.5 C4-.5 N N C7-.5 to C8-2. C6-.2 C8-.8 C4-.1 C4-.2 C5-.2 Top View S S N C8-.2 C4-.1 C7-2. C6-.1 C4-.2 C6-.2 C3-.3 to C6-.6 C6-.8 C4-.5 to C8-3. C5-.2 C4-.2 C4-.2 C4-.1 West Elevation C4-.1 C5-.1 S Figure 4-1: Sketch of cracks found in the specimen 9

4 Flexural cracking in the flange slab The first flexural crack in the top surface of the beam was observed running across the beam column intersection at a nominal specimen drift ratio of 1. %. The width of the crack at this stage was very small and varied from around.1 mm to.25 mm along the length of the crack. As the specimen drift was increased, flexural cracking propagated away from the beam column connection and cracks formed in earlier cycles widened. All the cracks formed were almost perpendicular to the beam spanning direction. Cracking in the longitudinal direction was not seen, indicating the non-existence of secondary effects. Crushing or spalling of concrete in flange slab was not observed. A sudden widening of one of the cracks was observed at the drift level of 3.5 %. A width of 2. mm was measured for this crack, which was significantly higher than the widths of other cracks at this stage. This main crack formed on top of the flange slab extended across the full width of the slab and its location coincided with the location of the main top longitudinal reinforcement curtailment point. During the last cycle with a nominal drift ratio of 4 %, the same crack widened to 4-5 mm and propagated in to the beam as shown in Figure 4-2. It was noted that at this instance a snapping sound came as a result of breaking internal mesh reinforcement. The main crack in the flange slab top surface, at the end of the test, is shown in Figure 4-3. As seen from Figure 4-2, the depth of the crack had extended over half of the depth. It was also noted that during the last cycle (75 mm displacement) of the test, a snapping sound came as a result of breaking internal mesh reinforcement. This was confirmed later during the rectification process, while removing loose material at the crack interface. This indicates that the mesh reinforcement has 91

5 reached its ultimate strength at 4% drift level. This will be further investigated in FEM analysis described in chapter 5. Main Crack Figure 4-2: Location and extent of main cracking after last cycle (North side beam) 92

6 Beam Spanning direction Figure 4-3: Main crack in flange slab top surface (North side beam) Flexural cracking in the ribbed beam The first flexural cracking in the beam surface was observed at a nominal specimen drift ratio of 1.2 %. The width of the crack at this stage was very small and it was around.1 mm. As expected the first crack formed at the bottom of the rib beam column connection. As the specimen drift was increased, these flexural cracks propagated away from the beam column connection. The type of cracking observed in the rib beam was quite normal. However, it was observed that the extent of cracking along the length of the beam had spread nearly 3/4 length of the span. Widths of these cracks were uniform compared with cracks found on the top surface of slab. 93

7 The bottom flexural crack at beam column interface gradually extended and joined the existing crack in the bottom level of the flange slab. As illustrated in Figure 4-4, some concrete crushing and spalling was seen at the rib beam-column interface, as a result of opening and closing of the crack. The concrete crushing near rib beam-column interface was started at a drift level of 3. %. Concrete crushing & spalling Figure 4-4: Concrete spalling at beam-column interface Flexural cracking in columns Figure 4-1 shows the location and extent of cracking in upper column. The first crack appeared at a drift level of about 1.2 %. The width of the crack at this stage was very small, around.1 mm. The crack extended across the full face of the column and had a depth of about 3 mm when originally observed. As the specimen drift level increased, the crack width also increased. However, there were no new cracks formed even at higher drift levels. 94

8 There was no cracking in the lower column. However, diagonal cracking was observed within the beam-column joint region as shown in Figure 4-4. Two of these cracks extended in to the lower column at a drift level of about 3. %. 95

9 4.2.2 Measured behaviour Hysteretic response The subassembly was tested to a maximum of 4 % nominal drift ratio. The hysteretic response of subassemblage was plotted as actuator load versus actuator displacement. The recorded response requires two types of corrections. The corrections are related to the reaction frame and the first correction was made to account for the movement due to the flexibility of the reaction frame. Flexibility of the frame was drastically reduced by the use of cross bracings. The recorded hysteretic response could be corrected for the above effect by recording the hysteretic response of the reaction frame. However, this measurement was not taken during the test and the correction was done using the hysteretic relation obtained from a mathematical relation developed by Stehle (22) for a similar frame. The second type of correction to be made was to account for the slip in bolts and pins. The pins at the top and bottom of the column were required to join the column to the test frame and actuator respectively. Since there was a clearance of about 2 mm between the pin and the hole, a pin slip was expected. A horizontal discontinuity was observed in the hysteretic results during the reversal in loading direction as can be seen in Figure 4-5. A slip correction of 2 mm was applied at this point to smoothen the curve as the loading direction changed (Figure 4-5). The fully corrected hysteretic response of the subassemblage is shown in Figure 4-6. The equivalent full-scale hysteretic response can be obtained by multiplying the recorded load by a factor of 4 and the recorded displacement by a factor of 2. 96

10 Actuator Load (kn) 2 mm slip Displacement (mm) 2 mm slip Figure 4-5: Hysteretic response showing pin slip in subassemblage Load (kn).8-1.2% Drift 2.-4.% Drift Displacement (mm) Figure 4-6: Fully corrected hysteretic response 97

11 The test subassembly exhibited good displacement capacity with no degradation of overall lateral strength up to 3 % drift. It was observed that in the positive load cycles the maximum actuator load was achieved in cycle 8 with 3 % drift ratio. However, in negative load cycles it was achieved in cycle 6 with 2.% drift ratio. The secant stiffnesses of the specimen at the end of the test and at 3% drift were 29 % and 56 % of the original respectively. Degradation of stiffness was mainly associated with flexural cracking, reinforcement yielding and slippage of beam column bars through the joint. As can be seen from the fatness of the hysteretic response (Figure 4-6), the system as a whole seems to have little energy absorption capacity Strain gauge readings Strain gauge readings were recorded continuously during the test, except when the actuator loading was paused, such as at peak of each load cycle for checking of cracks, etc. Data were recorded on a computer via the data logger. The selected strains versus load plots are presented within the body of the thesis for discussion purposes. Other strain plots are given in Appendix-E. Beam top longitudinal reinforcement Figure 4-7 shows the stain history of a top main reinforcement bar at the north column face. As can be seen, the yielding occurs at a load of 75 kn, during the cycle 6 (2. % drift). The strain increases under positive moment to 3 microstrain at the maximum 4% nominal drift. Under the negative moment loading, a tensile strain of 4 microstrain is attained at 4% drift level. At low drift, no slip was observed and small cyclic tension and compression is recorded. However, in cycle 3 (1. % drift), strain did not increase in compression when the load reversed direction, rather remained constant in compression 98

12 region, indicating the onset of bond deterioration. The occurrence of tensile strain under a positive moment (for beam top reinforcement) also indicates the presence of bar slip. Gauge BTG Strain ( micostrain ) 1 BTG4 8 6 Yield Strain Actuator Load ( kn ) Figure 4-7: Strain history of a top beam bar at north column face (East corner) The strain history of other top main reinforcement on the north column face is shown in Figure 4-8. The reinforcement yielding occurs during cycle 6 (2. % drift) as per the previous strain plot. During the same cycle (cycle 6), strain was increased from 2 to 13 microstrain. Under the negative moment loading, a maximum tensile strain of 137 microstrain was attained at 2.5% drift level (cycle 7). The tensile strain was gradually reduced at subsequent cycles and finally at cycle 9 (4 % drifts level) the strain was reduced to 12 microstrain. The maximum strain values observed in this strain gauge (BTG2) was about three times higher than the maximum recorded in gauge BTG4. One would expect the same strain history from both gauges, as they were located symmetrically on the same column face. The difference in strain behaviour depicted in 99

13 Figure 4-7 and Figure 4-8 could be due to number of reasons such as: (1) Proximity of strain gauge to cracks leading to a higher strain value. (2) Slip of bar and hence relief of strain (3) The strain gauge located at the local yielding region of the bar, hence higher strain gauge reading. This difference in strain behaviour was observed from the test, as cracking on flange top surface near gauge BTG2 was higher than the other half of north beam. Gauge BTG Strain ( micostrain ) BTG2 Yield Strain Actuator Load ( kn ) Figure 4-8: Strain history of a top beam bar at north column face (West corner) Beam bottom longitudinal reinforcement Figure 4-9 shows the strain history of the bottom main reinforcement at the north side beam. The maximum strain recorded was well below the yield strain. A similar strain plot was obtained for south side beam bottom bar as well, indicating that stain gauges are not in a critical region. Both plots show the gradual increase in strain with increasing drift ratio. 1

14 Strain (micostrain) Gauge BBG Yield Strain BBG Actuator Load ( kn ) Figure 4-9: Strain history of the beam bottom main bar (North side) Column strains Figure 4-1 shows the strain history of strain gauge in a southeast corner (SEC2) bottom column bar. A tensile strain of 1 microstrain, which is well below the yield, was developed at the maximum 4 % nominal drift. Southwest corner top column strain gauge (SWC1) recorded the maximum strain reading of 18 microstrain. Similar low and high maximum strain readings were observed for other bottom and top gauges respectively. This difference in strain behaviour was observed from the test, as slight cracking on upper column was seen while no cracking was observed on the lower column. 11

15 Strain ( micostrain ) Gauge SEC SEC2 Yield Strain Actuator Load ( kn ) Figure 4-1: Strain history of strain gauge in a southeast corner bottom column bar Displacement transducer readings As described in chapter 3, displacement transducers located on the specimen close to the joint were used to calculate the curvature of the section in regions of expected plastic hinging. Figure 4-11 and Figure 4-12 show the moment curvature plots of north and south beam respectively. Curvature of each beam was calculated using the readings obtained from transducers located in top and bottom of each beam. The bending moments of beam at the column face were calculated based on the recorded actuator force. These north and south beam moment curvature plot show a significant difference in behaviour. This may be due to the fact that the transducer readings are measured over a finite length. Many cracks can occur within that distance, giving an average value for that finite length. The cracking observed on either side of the column was not symmetrical during the test, hence a significant variation of curvature can be expected. 12

16 1 75 Bending moment (knm) Beam curvature (rad/m) -1 Figure 4-11: Bending moment versus beam curvature (North) Bending moment (knm) Beam curvature (rad/m) Figure 4-12: Bending moment versus beam curvature (South) Load cell values Figure 4-13 shows the variation of column axial force during the test. As mentioned earlier in chapter this prestressing force was applied using four prestressing strands. The prestressing force in strands remained constant at approximately 4 kn. The force only 13

17 varied by as little as 4%. Therefore, it is reasonable to assume that the column compression force remained constant at 4 kn during the test. 45 Column axial load (kn) Actuator Load (kn) Figure 4-13: Column prestressing force versus Actuator load Performance assessment Strength behaviour As seen from Figure 4-6, first subassembly test exhibited good displacement capacity with no degradation of overall lateral strength up to 3 % drift limit. It was observed that in the positive load cycles the maximum actuator load was achieved in cycle 8 with 3 % drift ratio. However, in negative load cycles it was achieved in cycle 6 with 2.% drift ratio. As seen from Figure 4-11 and Figure 4-12, both north and south beams reached the maximum bending negative and positive moment at cycle 8 (drift 3.%) and cycle 6 (drift 2.%) respectively. The attained strengths of various components of the subassemblage were calculated from the hysteretic response. The theoretical capacities of subassembly members were 14

18 calculated using the measured material properties. The experimentally attained actions are compared to the theoretical capacities and are given in Table 4-1. The experimentally attained column moments were determined by multiplying the column shear with the column height. The shear force at the top and bottom columns was same as actuator force. It should be noted that the columns did not yield and did not reach their ultimate capacities. The beam moments were determined by moment-curvature hysteretic response shown in Figure 4-11 and Figure The beams were yielded in both positive and negative directions. Both the positive and negative beam moments reached approximately 115% and 6% of capacity. The lower negative moment reached is due to the premature failure at the main top reinforcement curtailment point. The shear value attained was much below the capacity, mainly due to the fact that the rib beam has a very high shear capacity. 15

19 Table 4-1: Comparison of attained actions and theoretical capacities Design parameter Units Experimentally attained actions Actual theoretical capacity Experimental / Theoretical Bottom column 2 -Moment Capacity knm Shear capacity kn Top column 3 -Moment Capacity knm Shear capacity kn Beam (at column face) -Negative moment capacity knm Positive moment capacity knm Negative shear capacity kn Positive shear capacity kn Stiffness behaviour The degradation of stiffness can be seen from the applied actuator load versus displacement plot (Figure 4-6). The average stiffness of the specimen is calculated by 2 Column axial load taken as 4 kn- for calculations see Appendix-C 3 Column axial load taken as 345 kn- for calculations see Appendix-C 16

20 finding the slope of the line joining peak-to-peak points of the hysteretic loops. This method was recommended by Durrani and Wight (1985). The specimen experienced loss of stiffness as the drift ratio increased. This is due to the concrete cracking, yielding of reinforcement and the pull out of the beam longitudinal reinforcement from the joint. Comparison of stiffness degradation The stiffness degradation of the Corcon subassemblage was compared with the results of a similar subassembly test reported by Durrani and Wight (1987). This test series consisted of three interior beam-to-column sub-assemblages to study the effect of the presence of a floor slab on the behaviour of beam-column connections during an earthquake. The overall size and cross sectional details of all three specimens were same. A typical column height of 2248 mm and beam length of 2496 mm were used in all specimens. The typical member cross sections were; Main T Beam 419x279 mm with 1 mm thick and 13 mm wide slab, Transverse Beam-381x279 mm, column-362x362 mm. The three specimens were tested with different joint shear stress and the different amount of joint transverse reinforcement. The design of frame members was based on the ACI Building code. The length of the beams and the height of the columns represented one half of the span and the storey height, respectively, which is exactly similar to Corcon test specimen. The hysteresis loops of both systems were used to determine the stiffness degradation. The average peak-to-peak stiffness degradation of the specimens is illustrated in Figure For each specimen the stiffness is shown as a percentage of the initial stiffness. As reported by Durrani and Wight (1985), different levels of joint shear stress and joint confinement reinforcement have very little effect on the stiffness degradation of the specimen. They noted that the loss of average peak-to-peak stiffness at the end of the 17

21 seventh cycle (4% drift) was approximately the same magnitude for all three specimens in spite of the different level of confinement and the joint shear stress. Corcon beam showed sudden drop in stiffness at its last cycle (9 th cycle), which is clearly due to the main crack development near the top reinforcement curtailment point at the end of cycle 8 (3.% drift). The average peak-peak stiffness of the Corcon specimen at the end of the last cycle (4% drift) and at 3% drift was 29 % and 56 % of the original stiffness respectively. However, Corcon specimen shows comparatively good performance up to the 3. % drift level. 12 Stiffness Degradation (%) Specimen 1 (D&W) Specimen 2 (D&W) Specimen 3 (D&W) Corcon S Rel. Storey Drift (%) Figure 4-14: Stiffness degradation of Corcon and other specimens Energy dissipation The hysteretic response of the specimen provides a measure of the energy dissipated by the subassemblage during the test. The energy dissipated through damage in the specimen during a particular cycle is equivalent to the area enclosed by the corresponding loop. For this specimen, the hysteretic loops were thin which indicate low level of energy 18

22 dissipation. The energy dissipated by the specimen and the equivalent viscous damping ratio (h eq ) for the loading cycles is presented in Table 4-2. As described earlier in chapter 3.4.2, the recorded nominal drift ratio was adjusted to account for pin slip at connections. The corrected hysteretic response was used to calculate the viscous damping. Table 4-2: Energy dissipation and equivalent damping ratio Cycle Nominal Corrected Energy dissipated Equivalent number drift ratio drift ratio in half cycle viscous damping (%) (%) (joules) ratio (%) At.4 % drift, frame slip interferes too much so that a reliable calculation cannot be made. 19

23 A desirable behaviour for a beam-column subassemblage under cyclic loading implies a sufficient amount of energy dissipation without a substantial loss of strength and stiffness. As can be seen from Table 4-2, there is a gradual increase in h eq, indicating that higher level of energy being dissipated as the drift level increased. The sudden increase in h eq in cycle 6 (2% drift) was due to the first yielding of beam reinforcement. The above behaviour is clearly seen from the drift ratio versus equivalent damping ratio plot shown in Figure Generally, the equivalent viscous damping ratio corresponds well with first yielding of reinforcement, repeated yielding, initiation of cracks and widening existing cracks. A maximum equivalent damping ratio of 1.4% is calculated. This value was obtained at cycle 8 (3% drift) and at the last cycle, equivalent viscous damping ratio was again reduced. This may represent the severe loss of strength after the main cracking observed at the end of cycle Equivalent viscous damping ratio for cycle (%) Corrected specimen drift ratio (%) Figure 4-15: Drift ratio versus equivalent viscous damping ratio 11

24 Ductility and displacement capacity The displacement ductility factor of the specimen may be determined using the method presented in chapter 2, as shown below: The displacement ductility max, y Where max = Maximum displacement, y = yield displacement Since the hysteretic response of the structural components may not have a well-defined yield point, it is usually difficult to determine the displacement at yield. However, it was revealed from various plots such as hysteretic response of strain gauges and hysteretic response of north and south beam moment curvature, that yielding started at cycle 6 (2% drift). The ultimate maximum displacement of the subassemblage s hysteretic response could be taken as the maximum applied displacement at last cycle (4% drift). The displacement corresponding to this drift level could be taken as the subassemblage s maximum displacement due to severe cracking and high strength degradation at this level. Hence, a maximum displacement ductility ratio equal to 2 (ratio between 4% drift and 2% drift) can be calculated nd interior specimen Observed behaviour General As described earlier in chapter 3.7, the first test specimen was retrofitted with CFRP. This involved a major rectification of the specimen with up to.3 mm wide cracks repaired using epoxy injection in addition to CFRP. 111

25 The specimen appeared to perform very well under lateral deformations up to 4% drift level. The damage that was observed appeared to be very much moderate compared to the first specimen. As for the first specimen, there were no sign of cracks due to secondary effects such as torsion in the specimen Types and formation of cracks All cracks were numbered according to the cycle number in which they were first seen, as for the first specimen. However, there is a greater possibility that some of the cracks may have formed in an earlier cycle but were too fine to be detected. This issue is more critical in this test specimen, as the large amount of top flange and rib beam in the joint area were covered with CFRP. However, all the possible cracks were recorded as in the previous specimen. Cracks formed under positive loading were marked in red, while those formed under negative loading were marked in blue. Figure 4-16 illustrates the cracks found in the specimen after completing all loading cycles. This gives a clear picture of the overall cracking pattern of the specimen. It should be noted that only cracks that are visible out side the central joint area is shown, as cracking in the cental portion of the rib beam was covered with CFRP. 112

26 C4-.1 C7.2 C4-.1 C4 -.2 C7.3 S C C6-.1 N East Elevation N C9 3. C7.3 Top View S N C7-.2 C6-.2 C6.3 C9-3. C6.2 C6.3 C7.3 C4-.2 C4-.2 W est Elevation C4-.1 C6-.1 S Figure 4-16: Sketch of cracks found in the repaired specimen 113

27 Flexural cracking in the flange slab. The first flexural crack in the top surface of the beam was observed running across the beam column intersection at a nominal specimen drift ratio of 1.6 %. The width of the crack at this stage was very small and varied from around.1 to.2 mm along the length of the crack. All the cracks formed were almost perpendicular to the beam spanning direction as in the previous test. The cracking in the longitudinal direction, crushing or spalling of concrete in flange slab were not seen. A sudden widening of cracks was not observed as in the first specimen. The number of cracks in the top flange at the end of the last cycle was considerably low compared to the damaged observed in the first specimen. The gradual widening of cracks with increased drift levels was observed. This was a significant difference when compared to the performance of the first specimen. The main crack observed at the drift level of 3. % was less than 1 mm in width compared to more than 2 mm in the first specimen. Other cracks formed at this stage were relatively very small. This main crack formed on top of the flange slab extended across the full width of the slab as in the previous specimen and its location coincided with the location of the previous main crack. The cracking may have occurred at the same location due to the broken existing reinforcement and termination of main bars within the slab. During the last cycle with the nominal drift ratio of 4 %, the same crack widened to 3 mm and closed during the loading reversal. The main crack in the flange slab top surface at the end of the test is shown in Figure As seen from Figure 4-17, the main crack was limited only to the flange slab depth and had not penetrated in to the web area as happened in the first specimen. 114

28 Main Crack Rectification done for the 1 st specimen main crack. Figure 4-17: Location and extent of main cracking after last cycle (North side beam) Flexural cracking in the ribbed beam. The first flexural cracking in the beam surface was observed at a nominal specimen drift ratio of 1. %. The first crack formed at the bottom of the rib beam column interface similar to the first specimen. As the specimen drift was increased, flexural cracking propagated away from the beam column interface. However, the number of cracks observed was less compared to the first test. As shown in Figure 4-18, the bolted steel plates at bottom of the rib beam to prevent CFRP delamination were seen bending outward due to the outward force generated by CFRP. Due to the inadequacy of steel plate provided, the CFRP layer near the joint area was delaminated from the built up chamfer as the drift level increased and finally at the drift level of 4% a severe cracking was observed as shown in Figure This was mainly resulted due to the outward force from the CFRP and the high compressive force in the rib beam, near beam column interface. 115

29 Outward bending of steel plate Figure 4-18: Part of rib beam (north side) Cracking above the chamfer area Figure 4-19: Cracking near the built up chamfer area 116

30 Flexural cracking in columns The upper and lower columns of the subassembly were not rectified, as cracks on these columns were less than.3mm. It was observed during the second test that same cracks that formed earlier widened and no new cracks formed. However, the cracks observed in the top column were wider than the first test and were around.25 mm. There was no cracking in the lower column as in the first test Measured behaviour Hysteretic response The hysteretic response of the second test was recorded as was done for the first test. The recorded response required two types of corrections similar to the first test. Displacement corrections were done following similar procedure as in the first test (see chapter ). The fully corrected hysteretic response of the subassemblage is shown in Figure 4-2. The subassembly was tested to a maximum nominal drift ratio of 4 %. It should be noted that last cycle (4.% drift) in the negative loading direction was not done due to the inadequate space between the external prestressing supporting steel frame and cross bracings. The equivalent full-scale hysteretic response can be obtained by multiplying the recorded load by a factor of 4 and the recorded displacement by a factor of

31 Load (kn) % Drift 2.-4.% Drift Displacement (mm) Figure 4-2: Fully corrected hysteretic response (second test) Similar to the first specimen, the corrected hysteresis response shows that the system as a whole has relatively low energy absorption. However, compared with the first specimen the second specimen has absorbed 25-35% higher energy (see the Table 4.2 and 4.4). Test subassembly exhibited good displacement capacity with no degradation of overall lateral strength up to the last cycle (4 % drift). A higher maximum strength than for the first specimen was attained due to the CFRP strips provided beyond the reinforcement curtailment point of the first specimen. It is clear that the adopted CFRP system has proven to be an effective technique to repair/strengthen the test specimen. Compared to the original, the retrofitted specimen has increased its lateral load resistance by 26% at peak load. At large displacements, the load was still increasing with positive stiffness, though with a stiffness lower than that of initial value. The degradation in stiffness is attributed to flexural cracking and loss of anchorage of both the beam reinforcement and the CFRP system. 118

32 Photogrammetry-based measurement As described in chapter 3, photogrammetry-based measurements enabled both global and local deformation of the test specimen to be followed during the testing. Figures 4-21 to 4-24 show the deformation of first row of photo-sensitive targets on the flange slab, for the load cycles from 1.2 % to 4. % drift. Figures 4-21 and 4-22 show the vertical and horizontal components of the movement of each target during the north side displacement of the actuator. Similar plots shown in Figures 4-23 and 4-24 give the vertical and horizontal movement of same target points during the south movement of the actuator. Deformation (mm) 6 1.2% Drift 1.6% Drift 4 2.% Drift 2.5% Drift 2 3.% Drift 4.% Drift Distance from column center (mm) Figure 4-21: Vertical deformation of the beam (North displacement of actuator) Max. Crack width = 2.26 mm Deformation (mm) Distance from column center (mm) 1.2% Drift 1.6% Drift 2.% Drift 2.5% Drift 3.% Drift 4.% Drift Figure 4-22: Horizontal deformation of the beam (North displacement of actuator) 119

33 The vertical deformation plots give a clear picture of the beam vertical deformation at the end of each load cycle and the horizontal deformation plots gives the axial deformation of the beam flange. The axial deformation along the beam flange slab could be used to locate the cracking in between photo-sensitive target points. As shown in Figure 4-22 the cracking during last cycle (4% drift) has increased to 2.26 mm. This matches perfectly with the crack width observed (3 mm) during the 4. % drift cycle. The slight variation of observed crack width and the value obtained from the graph, may be due to the change in location of measured point and the target location. These plots will be compared with the results obtained from finite element modeling in chapter 5. Deformation (mm) % Drift 1.6% Drift % Drift 2.5% Drift -5 3.% Drift -6 Distance from column center (mm) Figure 4-23: Vertical deformation of the beam (south displacement of actuator) Deformation (mm) % Drift 1.6% Drift % Drift 2.5% Drift -1 3.% Drift Distance from column center (mm) Figure 4-24: Horizontal deformation of the beam (South displacement of actuator) 12

34 Strain gauge readings on reinforcement Strain gauge readings were recorded in a similar manner as for the first specimen. The same strain plots presented for the first specimen was selected for the second test as well, in order to provide direct comparison. In addition, the results from selected strain gauges on CFRP are presented within the body of the thesis for discussion purposes. Beam top longitudinal reinforcement Figure 4-25 shows the strain history of a top main reinforcement bar at the north column face. The maximum strain recorded at the end of 3 % drift level was 14 microstrain, compared to the strain level reached in the first test of 39 microstrain. Similarly, in the second test reinforcement yielding occurred during the last cycle (4 % drift), whereas in the first test reinforcement yielding occurred during cycle 6 (2. % drift). It is very clear from the plot that the bar was strained to lower values than recorded strain in the first test. The reason for this is due to the contribution of well-anchored CFRP as flexural reinforcement in the negative moment area. 121

35 Gauge BTG4 5 BTG4 4 Strain (Micro Strain) 3 Yield Strain Actuator Load (kn) Figure 4-25: Strain history of a top beam bar at north column face (East corner) The strain history of other top main reinforcement on the north column face is shown in Figure The reinforcement yielding occurred during the last cycle (4. % drift) as per the previous strain plot. The maximum strain recorded at the end of 3 % drift level was 136 microstrain, which is consistent with the strain level observed in the other main top bar (14 microstrain). However, in the first test, the strains found were not compatible for the bars located symmetrically on the same column face. This could happen in an area where the cracking is very high. 122

36 Gauge BTG2 5 4 BTG2 Strain (Micro Strain) 3 Yield Strain Actuator Load (kn) Figure 4-26: Strain history of a top beam bar at north column face (West corner) Figures 4-27 and 4-28 show the strain history of gauges located at CFRP top strips at the north-east and north-west respectively. The maximum strain recorded at the end of 4 % drift level was 26 microstrain and 23 microstrain respectively, which are well below the designed maximum strain of 49 microstrain. It should be noted that sudden increase in strain observed in the strain gauges BTG2 and BTG4 (gauges located at top main reinforcement bars on north side) was not observed in CFRP strain gauges. Figures 4-29 and 4-3 show the strain history of CFRP top strips at 1 mm away from the column center on east and west side respectively. The very high strain increase observed in above strain plots are similar to the strain increases observed in BTG2 and BTG4 (gauges located at top main reinforcement bars on north side). The main cracking recorded in north beam coincide with the above strain gauge location. Therefore higher strain reading can be expected due to the proximity of the strain gauge to cracks. 123

37 Gauge CG2 5 Strain (Micro Strain) Actuator Load (kn) S N CG2 Figure 4-27: Strain history of top CFRP at north column face (East corner) Gauge CG1 5 4 S CG1 N Strain (Micro Strain) Actuator Load (kn) Figure 4-28: Strain history of top CFRP at north column face (West corner) 124

38 Gauge CG1 5 S CG1 N 4 Strain (Micro Strain) Actuator Load (kn) Figure 4-29: Strain history of a top CFRP at 1. m away from column (East side) Gauge CG9 CG9 5 S N 4 Strain (Micro Strain) Actuator Load (kn) Figure 4-3: Strain history of a top CFRP at 1. m away from column (West side) 125

39 Beam bottom longitudinal reinforcement Figure 4-31 shows the strain history of the bottom main reinforcement of the north side beam. The maximum strain recorded was well below the yield strain as in the first test, indicating that stain gauges reading is not influenced by the CFRP used in the bottom beam-column joint area for strengthening. A similar strain plot was obtained for south side beam bottom bar as well. Both plots show the gradual increase in strain with increasing drift ratio. It was noted that the strain increase was only 16 % for the 6% increase in drift level from 2.5 % to 4.%. Gauge BBG2 5 4 Strain (Micro Strain) BBG2 3 Yield Strain Actuator Load (kn) Figure 4-31: Strain history of the beam bottom main bar (North side) Figures 4-32 and 4-33 shows the strain history of CFRP north beam bottom strips on west side at 6 mm and 2 mm away from the column center respectively. The strain recorded in CFRP was in the same order as in reinforcement. The gauge CG13 (Figure 4-32) shows good bond behaviour during cyclic loading. However, gauge CG14 (Figure 126

40 4-33) shows very poor bond behaviour after 2.5 % drift level. This is due to the delamination of CFRP strips on both sides of the beam near the beam-column joint. Gauge CG13 15 Strain (Micro Strain) 1 S N CG Actuator Load (kn) Figure 4-32: Strain history of north beam bottom CFRP at 6 mm away from column (West side) 127

41 Gauge C14 15 Strain (Micro Strain) 1 5 S N CG Actuator Load (kn) Figure 4-33: Strain history of north beam bottom CFRP at 2 mm away from column (West side) Column strains Figure 4-34 shows the strain history of gauge SEC2 located in a southeast corner bottom column bar. A tensile strain of 15 microstrain, which is well below the yield, was developed at the maximum 4 % nominal drift. Figure 4-35 shows the strain history of SWC1 located in a Southwest corner column top bar. Gauge SWC1 recorded a maximum strain reading of 23 microstrain compared to strain observed in the first test of 18 microstrain. Similar low and high maximum strain readings were observed for other bottom and top gauges respectively. This difference in strain behaviour was observed in the first test as well, as slight cracking on upper column was seen while no cracking was observed on the lower column. 128

42 Gauge SEC2 5 4 Strain (Micro Strain) 3 Yield Strain 2 SEC Actuator Load (kn) Figure 4-34: Strain history of strain gauge in a southeast corner- bottom column bar Gauge SWC1 5 4 Strain (Micro Strain) Yield Strain SWC Actuator Load (kn) Figure 4-35: Strain history of strain gauge in a southwest corner -top column bar 129

43 Displacement transducer readings As for the first specimen, displacement transducers were used to measure beam curvature near the joint. Figures 4-37 and 4-38 show the moment curvature plots of north and south beam respectively. The bending moment and curvature of each beam were calculated following the same procedure used in the first test (chapter ). There is no significant difference between north and south beam moment curvature plots compared to the large difference observed in the first test. This may be due to the less cracking observed in the second test compared to the first test in the beam-column interface area. Both moment-curvature plots do not show any degradation of moment capacity Bending Moment (knm) Beam Curvature (rad/m) Figure 4-36: Bending moment versus beam curvature (North) 13

44 125 1 Bending Moment (knm) Beam Curvature (rad/m) Figure 4-37: Bending moment versus beam curvature (South) Load cell values Figure 4-38 shows the total column axial force variation during the test. It can be seen from the plot that the prestressing force in strands remained fairly constant at approximately 4 kn. The axial force varied by 7% compared with the 4 % variation observed in the first test. 131

45 Column axial load (kn) Actuator Load (kn) Figure 4-38: Column prestressing force versus Actuator load Performance assessment Strength behaviour As observed in Figure 4-37, first subassembly test exhibited good displacement capacity with no degradation of overall lateral strength up to 4 % drift limit. The attained strengths of various components of the second subassemblage are calculated from the hysteretic response as for the first specimen. The theoretical capacities of beams were calculated using the CFRP strips used to retrofit the specimen and ignoring the contribution due to reinforcement. The experimentally attained actions are compared to the theoretical capacities in Table 4-3. The experimentally attained column moments were determined by multiplying the column shear with the column length similar to the first test. The shear force at the top and bottom columns was same as actuator force. It should be noted that the upper column just 132

46 yielded and lower column did not yield. This is due to the shorter clear height of the lower column. The beam moments were calculated using the reactions in the vertical links at the beamends. The beams were yielded in both positive and negative moments. Both the positive and negative beam moments reached approximately 6% and 5% of capacity. The lower negative and positive moment reached is due to the fact that the test was carried out only up to the maximum lateral displacement capacity ( 75 mm. i.e. 4% drift) of the actuator. It should be noted that, even at 4 % drift level the load was still increasing at a lower rate than that of lower drift values. Similar to the first test, the shear value attained was much below the capacity, mainly due to the fact that the rib beam has a very high shear capacity. 133

47 Table 4-3: Comparison of attained actions and theoretical capacities (2 nd Test) Design parameter Units Experimentally attained actions Actual theoretical capacity Experimental / Theoretical Bottom column 5 -Moment Capacity knm Shear capacity kn Top column 6 -Moment Capacity knm Shear capacity kn Beam (at column face) -Negative moment capacity knm Positive moment capacity knm Negative shear capacity kn Positive shear capacity kn Stiffness behaviour The degradation of stiffness can be seen from the applied actuator load versus displacement plots (Figure 4-2). The average stiffness of the second test was calculated 5 Column axial load taken as 4 kn 6 Column axial load taken as 345 kn 134

48 similar to the first test. The specimen experienced loss of stiffness as drift ratio increased. However, the rate of stiffness degradation was low compared to the first test. This is due to the less cracking observed in the second test. Other main reason for above behaviour is due to both reinforcement and CFRP not being stressed to their yielding level. The second test specimen did not show sudden drop in stiffness up to the last cycle (4 % drift), rather stiffness degradation was gradual. This improved behaviour was due to the provision of top CFRP strips well beyond the original reinforcement curtailment point, thus avoiding excessive cracking. The average peak-peak stiffness of the second test specimen at the end of the last cycle (4% drift) and at 3% drift was 87 % and 78 % of the original stiffness respectively Energy dissipation The energy dissipation of the second test specimen was calculated similar to the first test specimen. As for the first specimen, thin hysteretic loops were observed, indicating the low level of energy absorption. The energy dissipated by the specimen and the equivalent viscous damping ratio (h eq ) for the loading cycles are presented in Table 4-4. Compared with the equivalent viscous damping ratio values obtained for the first test and the second test shows relatively lower values, indicating that the overall damage is less than the first specimen. 135

49 Table 4-4: Energy dissipation and equivalent damping ratio (2 nd Test) Cycle Nominal Corrected Energy dissipated Equivalent viscous number drift ratio drift ratio in half cycle damping ratio (%) (%) (%) (joules) Ductility and displacement capacity As explained previously for the first specimen (chapter ), it is usually difficult to determine the displacement at yield. The main top reinforcement did not yield until a nominal drift of 4 %. The second specimen did not reach to the ultimate displacement, as the test had to be terminated at 4% drift when the actuator reached its maximum capacity. At 4% drift level, the specimen exhibited no strength degradation. In fact, the strength of the specimen was still increasing. Hence, the ductility or maximum displacement cannot be expressed as in the first specimen. 7 At.4 % drift, frame slip interferes too much so that a reliable calculation cannot be made. 136

50 Durrani, A. J. and J. K. Wight (1985). "Behavior of Interior Beam-to-Column Connections Under Earthquake-Type Loading." ACI Structural Journal 82(3): Durrani, A. J. and J. K. Wight (1987). "Earthquake Resistance of Reinforced Concrete Interior Connections Including a Floor slab." ACI Structural Journal: pp Stehle, J. S. (22). The Seismic Performance of Reinforced Concrete Wide Band Beam Frames: Interior Connections, The University of Melbourne, Australia. 137