REINFORCED CONCRETE WALL BOUNDARY ELEMENT PRISM TESTING

HOLLY A J WRIGHT

Department of Civil and Environmental Engineering, University of Auckland, Auckland, New Zealand

Disclaimer– This conference paper has been submitted as partial fulfilment for the project requirement for the BE(Hon) degree. Although they have been assessed, no errors or factual information have been corrected or checked. ABSTRACT The purpose of this research was to observe the effects of confining reinforcement detailing on the ductility of well-detailed end regions of reinforced concrete walls, which are often termed “boundary elements.” Six boundary element prisms were tested under monotonic compression. The confining reinforcement detailing consisted of different combinations of hoops and ties, cover concrete thickness and transverse hoop spacing. All specimens were compliant with NZS3101:2006. The most ductile specimen achieved a full-height strain of 0.021, with a greater ductility spread over the specimen height. The other specimens had equivalent failure strains of 0.014, and exhibited much lower ductility spread over the specimen height. It was concluded that hoop unbending contributed to lower ductility in the specimens, and enabled out-of-plane translation of the gross section during failure. The use of single piece ties (uni-ties) led to a response with ductility greater than or equal to that of traditional detailing, as the single piece ties had less susceptibility to hoop unbending. Also, reducing concrete cover led to an improved ductility spread. A critical s/db ratio where hoop unbending will initiate for uni-ties was determined to be between 2.5 and 3.75. INTRODUCTION Reinforced concrete (RC) walls are effective in buildings as a means of resisting lateral seismic loading. RC walls consist of two regions: the web, and the boundary elements (BEs) ( Figure 1). Under lateral loading, the boundary elements are subjected to cyclic tension and compression loads, with higher strain demands comparatively to the web. Therefore, care must be taken in BE detailing, so that under the compression phase of the cycle they can behave in a ductile manner (i.e., withstand large strains prior to failure).

Figure 1: Plan view of a reinforced concrete wall, indicating web and BE locations

Web Boundary Element Boundary Element

The New Zealand Concrete Standard, NZS3101:2006 includes requirements aiming to achieve ductile responses of BEs (Standards New Zealand, 2006). Despite this, recent testing of code compliant boundary element specimens has suggested that BEs are failing before the expected degree of ductility is obtained. Three reasons for this have been identified. Firstly, hoop unbending in specimens results in effectively unrestrained longitudinal bars. This leads to premature bar buckling and core crushing, reducing the spread of ductility (Welt, 2015). Secondly, thick cover concrete results in concentration of plastic deformation, causing little spread of ductility and low ductile failure (Arteta, 2015). Thirdly, large spacing between transverse reinforcement allows longitudinal bars to buckle more easily, resulting in failure (Arteta, 2015; Massone et al., 2014). The objective of the research discussed in this report was to investigate factors that are known to give low ductility responses in boundary elements. The objectives are as follows:

• Compare the ductility of BEs detailed with traditional confining reinforcement, consisting of hoops and 180º-180º ties, with that of BEs confined using single piece ties (uni-ties).

• Observe the effects of changing the cover concrete thickness on the ductility of BEs.

• Compare the effects of different transverse reinforcing hoop spacings on ductility.

BACKGROUND Current requirements of the New Zealand standard for concrete structures – NZS3101:2006 The New Zealand Standard for Concrete Structures provides recommendations for confinement that aim to produce ductile responses for BEs. NZS3101:2006 Clause 11.4.6.3, specifies that longitudinal bars at the perimeter of the section require restraint with a hook or tie, unless the adjacent longitudinal bars are spaced at 200mm or less, in which case restraints can be placed on alternate bars. Where longitudinal bars are restrained by hooks, the bend angle shall be at least 135º. Where a longitudinal bar is restrained in the corner of a hoop, the angle shall be 90º. Transverse hoop spacing in ductile plastic regions shall be no more than six times the longitudinal bar diameter. In limited ductile regions, transverse hoop spacing should not be greater than ten times the longitudinal bar diameter. Past testing on boundary element specimens Monotonic compression tests on boundary element prism specimens were carried out by Arteta (2015), Welt (2015), Massone et al. (2014), and Mander et al. (1988a). The methodology, results and conclusions of these are discussed below.

Specimen detailing

The main variations between the test specimens in the four test programmes involved transverse reinforcement arrangements, varying between rectilinear configurations, and 180º, 135º, and 90º crossties. Also, longitudinal and transverse bar diameters and quantities were differed, affecting the standardised longitudinal bar slenderness ratio, s/db (where s is the transverse reinforcement spacing and db is the longitudinal bar diameter). This ratio was varied to determine the tendency of longitudinal bars to buckle. Instrumentation Arteta (2015) applied load through a pressure capsule (PC) on top of the loading yoke. Potentiometric displacement transducers (PDT) measured the relative displacement of the top and bottom head and gave overall strains. Wirepots measured out-of-plane displacement. Strain gauges at 127mm intervals detected and measured localised strain concentrations.

Welt (2015) used a 4450kN test frame to apply load. String potentiometers measured absolute displacement. Strain gauges were installed on reinforcement bars prior to casting, and measured strains in longitudinal and transverse reinforcement. Photogrammetry was used to detect surface deformations, and expansion gauges at the top and bottom measured lateral sample expansion. Mander et al. (1988a) used a 10MN servohydraulically controlled actuator to apply load. Four linear potentiometric displacement transducers (LPDTs) recorded longitudinal strains over the BE’s central section on each face. Strain gauges measured transverse strains in the transverse reinforcement across the BE’s central section. Massone et al. (2014) used external linear variable differential transformers (LDVT) to measure longitudinal strains over the specimen height. The LDVTs were attached away from the specimen face, so spalling had no effect on readings. Four internal LDVTs measured average longitudinal strain over the sample’s central portion, and were only accurate pre-spalling. Reinforcement bending after this point deemed measurements inaccurate. Results Mander tested eight BE specimens under monotonic compression loading. The results and cross sections of four specimens can be seen in Figure 2. Note that wall 1 underwent premature failure (due to premature wall buckling, as the walls were quite slender) and wall 2 (with similar detailing but a higher strain rate) was used in its place. Walls 2 and 4 were identically detailed except wall 2 used rectangular hoops and wall 4 used 180º-180º ties. These two walls performed similarly, suggesting rectangular confinement can be swapped with 180º-180º ties with little ductility loss. Wall 3, with 50mm transverse reinforcement spacing (hence a higher s/db ratio compared to other specimens which had spacings of 25mm) achieved less ductility, suggesting that the larger the s/db ratio, the less ductility achieved. Massone et al (2014) tested 11 BE specimens. The stress-strain behaviour of P5, P13, P17 and P18 is shown in Figure 3. P17 and P18 were detailed the same, except for the s/db ratios of 8.3 and 5.6 respectively. P5 and P13 are the same as P17 and P18 respectively, however the yield strength of the transverse reinforcement was 330MPa rather than 495MPa. For both transverse reinforcement strengths, the specimen with the larger s/db ratio had the lower ductility.

Figure 2: Stress-Strain response for Walls 1/2, 3,

and 4 - (Mander et al., 1988a)

Figure 3: Stress-Strain response of P5, P13, 917

and P18 - (Massone et al., 2014)

Figure 4: Stress-strain response of W10,

W11, and W12 (Arteta, 2015)

Arteta (2015) tested thirteen BE specimens and kept the horizontal spacing constant for all tests. The vertical spacing of transverse reinforcement was changed, as was the transverse reinforcement type, which resulted in differences in performance.

Figure 4 shows a specimen performance comparison of W10, W11 and W12. W11 was used as a control test specimen. W12 had a larger transverse bar diameter than W11, and W10 had smaller longitudinal bars than W11. The performance of W10 and W11 was similar. W12 had a gentler descending slope than W11, showing greater ductility. These results suggest that transverse bar thickness has more of an effect on ductility than longitudinal bar thickness.

Welt (2015) carried out testing on five BE specimens under monotonic compression loading. Welt identified four damage states: spalling of cover concrete, core crushing, longitudinal bar buckling, and crosstie failure, as seen in the force-displacement graph in Figure 5.

CS7 was detailed with overlapping hoops and clearly had the best ductility. CS7 achieves all damage states at higher strains than other specimens, indicating the use of overlapping rectilinear hoops in the confinement arrangement is ideal. In reality however, this is a labour intensive option, and is unlikely to be used.

CS5 and CS9 had the same confinement arrangements, only differing in the greater transverse hoop spacing of CS9 (127mm compared to 64mm of CS5). CS9 had lower ductility than CS5, with buckling at lower strains. This indicates that larger s/db ratios lead to lower ductility.

Welt (2015) found that some specimens had 135º-135º ties which unbent (Figure 6). He observed that the specimens in which this occurred exhibited lower ductility than in those where it did not. Specimen testing carried out by Arteta (2015), Massone et al. (2014), and Welt (2015), all saw damage concentration at the failure zone, and consequently little spread of ductility over the specimen height. Conclusions from testing The following conclusions can be drawn from the above tests:

• Unbending observed in confining reinforcement leads to low ductility in boundary elements. (Welt, 2015)

• 135º-135º ties are more susceptible to unbending, decreasing ductility. Longitudinal bars restrained in corners of rectangular hoops gives higher ductility. (Welt, 2015)

• A greater s/db ratio reduces ductility, credited to increased buckling likelihood. (Arteta, 2015; Massone et al., 2014; Welt, 2015)

Figure 5: Stress-strain response of CS5,

CS7, CS9, CS11, and CS15 (Welt, 2015)

Figure 6: Unbending of confining reinforcement (Welt, 2015)

• The majority of specimens underwent localised damage over small regions, so the ductility spread was limited by deformation concentration in one area (Arteta, 2015; Massone et al., 2014; Welt, 2015). This was common in specimens with large concrete cover (Arteta, 2015).

• No transverse reinforcement fracture was observed in any of the tests discussed above (Arteta, 2015; Mander et al., 1988a; Massone et al., 2014; Welt, 2015).

TEST DESCRIPTION Specimen details To address the issues identified in the past testing discussed above, six reinforced concrete boundary element prisms were tested under monotonic compression. Test specimen details are shown in Table 1 and Table 2. Specimens were built with identical dimensions of 150mm (depth) x 300mm (width) x 1500mm (height). These represent the boundary elements at the bottom storey of a reinforced concrete wall at one-half-scale. The top and bottom 300mm of the specimens were confined using steel plates on all sides (Figure 7) to provide specimen end zones similar to footings, which were intended to prevent failure at the ends of the specimens. Specimens were therefore defined as the central 900mm between top and bottom plates. The test variables were as follows and are shown in Table 1 and Table 2:

• Concrete cover was either 10mm (BE2, BE4, BE5 and BE6) or 20mm (BE1 and BE3).

• The confining reinforcement was one of two arrangements. BE1 and BE2 had traditional hoops and ties, where the ties were 180º-180º. BE3, BE4, BE5, and BE6 had uni-ties, which were formed from a single length of R6 bar, bent to form the shape shown on Table 1. Uni-ties allow every longitudinal bar to be restrained in a corner.

• The spacing of the confining reinforcement was either 40mm, 60mm, or 80mm. BE1, BE2, BE3, and BE4 had the smallest transverse reinforcement spacing of 40mm. BE5 had a spacing of 60mm, and BE6 was spaced at 80mm.

Table 1: Cross section of test specimen detailing Table 2: Elevation view of test specimen detailing

Test design Test set-up The test rig was set up within close proximately to the strong wall, with the specimen centred on a 450mmx450mm steel plate between four holes in the strong floor. As part of each test set up, dental plaster was spread on the plate and the specimen was lowered onto it. The specimen was then manoeuvred until level in all directions. This was checked using a spirit level. Plaster was also spread on the top of the specimen and a 300mmx300mm plate was placed upon it and levelled. The loading beam, which consisted of two 530UB92.4 I-beams connected with welded tabs, was then lowered onto the specimen (Figure 7). Two steel frames that fit around the loading beam were post-tensioned to the strong wall to restrain the loading beam on both ends against lateral translation. Vertical load was applied at the centre of the loading beam and at both ends. The central configuration is made up of two spreader beams (C-sections joined with welded tabs) tack-welded to the loading beam. Load cells and 60T jacks were placed on the spreader beams, and 32mm Macalloy bars were threaded through this arrangement and through holes in the strong floor, where they were anchored using plates and nuts. The bars were also anchored at the top using washers and nuts. The setup on the ends of the beams followed a similar configuration, except the jacks were placed directly onto the beam with the load cell on top of it. The spreader beam sat on top of this and 28mm Macalloy bars extended down from their points of anchorage on the spreader beam, into the strong floor where plates anchored them underneath. Load was applied to the specimen by pumping the hydraulic pumps, connected to the jacks

by hoses. Pumping up the jacks put the Macalloy bars into tension which, due to the

anchorage under the floor, put the loading beam into compression, applying a compression

load to the boundary element prism.

As can be seen in Figure 7, the two wooden frames on either side of the specimen were used as catch frames in case the beam dropped suddenly upon failure. Instrumentation

Load cells were positioned above or below each jack, as seen in Figure 7, and were used to monitor the load in each jack during testing, and the data was recorded. Portal gauges measured local strains between the intervals indicated in Figure 8. This same configuration was used on both north and south faces of the specimen. The locations of wirepots are shown in Figure 8. Two wire pots extended down the length of the specimen, measuring total specimen displacements. The other eight wire pots were connected to brackets on the strong wall, and the strings extended to the specimen. These measured out of plane displacement.

Figure 7: Loading rig and test specimen set-up

A tilt metre was set up on the top of the loading beam in the centre. This measured tilt in the strong and weak axis, which was monitored throughout testing. All of the instrumentation mentioned above was connected via cables into the DAQ. The data was recorded using the CompactDAQ program. Photogrammetry was also carried out, with two DSLR cameras located directly in front of the specimen, taking photographs at 10 second intervals. Failure footage was recorded using GoPros within close proximately to the specimen. Loading protocol Under lateral seismic loading, boundary elements are subjected to cyclic tension and compression loading. There are three main failure modes that can occur under this loading:

• Concrete crushing under compression

• Longitudinal bar buckling resulting from tensile strains

• Global section buckling resulting from large tensile strains Previous testing by (CW Hilson et al., 2014a) has shown that the failure stress in both cyclic tension and compression loading and monotonic compression loading is similar. Therefore, it is acceptable to carry out this research using monotonic compression loading, and disregarding cyclic effects. Test procedure Throughout testing, the aim was to load the specimen uniformly, keeping the tilt in both directions at zero. Initially, the outside jacks were evenly loaded to start the loading process without causing any tilt. Once approximately 200kN was reached in each outside jack, the middle jacks would be pumped, loading two at a time, on the diagonal. Any sign of the beam tilting would prompt pumping of the corresponding jack to fix this tilt. This procedure was strictly followed to ensure the tilt remained very close to zero. Each boundary element was loaded until failure. The force in each jack was monitored throughout loading, to ensure the jacks or load cells were not exceeding their limits. Load cell readings were also checked with the pressure gauge readings on the hydraulic pumps as a safety precaution. Material property tests

Concrete cylinder tests

Concrete cylinders were tested on the same day that a specimen was tested. Three cylinders were tested for each specimen and the average value used as the concrete compressive strength. Table 3 shows the strengths recorded for each cylinder. It is clear that the cylinders tested with BE2 had lower strengths. From visual inspection of the cylinders and the

Side View

Front View

Legend

Figure 8: Instrumentation layout – wirepots and portal

gauges

Table 3: Results of concrete cylinder testing

(values in MPa)

results shown in Table 3, the BE2 cylinders were poorly prepared. Hence the concrete strength of BE2 will be considered as similar to that of the other specimens. No time effects were observed for the concrete strength, hence the average strength of all cylinders was suitable for calculation use. Steel tensile tests Tensile tests were performed on reinforcing steel samples and the results are shown in Table 4. The results show that the GR500 (D16H) steel was close to the expected values. The R6 values however were more variable. The second R6 bar in Table 4 has a yield strength of 100MPa. Upon inspection of the bar, it was in poor condition prior to testing, and failure could have initiated prematurely from a location of damage. The higher yield stresses of the remaining two R6 bars can be credited to strain hardening which will have occurred prior to carrying out these tests. R6 bar is coiled after production for storage, causing strain hardening. It was noted during testing that these bars did not have a well-defined yield plateau. Results As shown in Figure 9, there were two failure shapes observed in the six test specimens. The first of these shapes was an out of plane failure shape. This shape was observed with failure concentrated at the top in BE1, BE2, BE3, and BE6. BE5 also exhibited this out of plane shape at failure, however deformation was concentrated at the bottom. BE4 had a unique failure shape and underwent radial expansion in the failure zone at the top, with very little out of plane deformation. All specimens underwent one of two failure mechanisms: tie fracture only (BE4) with no hoop unbending, as shown in Figure 10, or a combination of hoop unbending and tie fracture (all other specimens), shown in Figure 11 and 12. The following sequence of events led to the failure mechanisms observed:

1. Concrete cover spalls 2. The concrete core cracks and attempts to expand outward 3. Confining reinforcement is engaged in tension 4. The corner hoops unbend and the stresses are transferred to the central ties (this step

did not occur in BE4) 5. Tie fracture 6. Concrete crushes and longitudinal bars buckle.

BE5 and BE6, which had larger spacing of transverse reinforcement than the other four specimens, saw little spalling prior to failure. Rather, these two specimens failed in more of an explosive manner, and the cover was lost all at once. BE1, BE2, BE3, and BE4 however, underwent spalling over a greater length of the specimen before failure occurred. These four specimens had lost substantially more cover by the time longitudinal bar buckling and core crushing occurred.

All failure occurred within the specimen height and did not extend into the sections confined by the steel plates at the top and bottom.

Table 4: Results of steel tensile testing

Figure 13 is the normalised stress-strain plot for the six boundary element specimens. The normalised force, P/P0, is the force applied to the specimen (P) divided by the theoretical compression strength P0 (Equation 1).

P0 = ( Ag - As ) * f’c +As * fy (1)

The normalised strain is the overall displacement divided by the clear height. From Figure 13 it is clear that BE1 sustains the highest strain of 0.027. BE4 has the second highest strain sustained, at 0.021, and all other specimens equally sustained the lowest strain of 0.014. Figure 14 shows the local strains over the specimen height. The locations of strain concentration in the failure bays in Figure 14 matches the failure shapes seen in the photographs in Figure 9, with the damaged section undergoing the most strain. As seen in Figure 14, BE1 undergoes the most local strain in the failure bay of 0.142, but drops down to only 6.5% of the maximum strain in the bay below. A similar response is observed in BE3, BE5 and BE6. Lower local peak strains are reached by BE4, however 33.6% of strain is spread into the next bay. BE2 follows a similar response to BE4, but with a slightly smaller spread. The local strains were determined by the displacement measured in the portal gauge, divided by the gauge length. The lateral deflections at failure are shown in Figure 15 and represent post-failure shapes of the specimens. BE4 shows little lateral deflection, due to the radial expansion failure shape. BE2 shows a negative lateral displacement during failure, correlating with the failure shape shown in Figure 9.

Figure 9: Failure shapes observed in the specimens

Figure 10: Tie

fracture only

Figure 11: Tie

fracture and hoop

unbending (uni-

ties)

Figure 12: Tie

fracture and hoop

unbending (hoops

and ties)

As the top and bottom of the specimen were ’fixed’ against lateral displacement, the middle of the specimen displaced in the negative direction, causing the out of plane failure shape. This phenomenon was observed in all specimens with an out of plane shape at failure.

DISCUSSION BE1 control issues During loading, the southern load cells were connected to a strain gauge card that had not been properly switched to a load cell card, therefore the gain was double what it should have been. Consequently, the readings of the southern load cells were twice that of the northern load cells. As this was the first test, it was thought that this was a loading issue and the northern jacks were loaded in an attempt to fix the force balance. Consequently, the specimen was subjected to a tilt of up to 1 degree. Following that, the hose connected to the north western jack began to leak, so this jack was not balanced with the north eastern jack, and the loading beam had a tilt of up to 2 degrees in the major axis. Consequently, the specimen was loaded on a strain gradient (Figure 16).

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Lateral Deflections of Specimens at Failure

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Figure 13: Normalised force versus total strain plot

Figure 14: Local strain verses height

plot

Figure 16: Non-uniform loading in

BE1 compared to the uniform loading

of the other specimens

Figure 15: Lateral deflections of the specimens at

failure plot

Seeing as BE1 was not loaded uniformly like the other specimens, the results cannot be directly compared to the other specimens and therefore will not be further discussed in this report. Specimen failure shape BE4 had a unique failure shape and exhibited no hoop unbending. Seeing as all other specimens had hoop unbending and an out of plane failure shape, this suggests that there is a correlation between the presence of hoop unbending and the failure shape achieved. Upon inspections of the failed specimens, it was apparent that the out of plane failure shape is attained from unbending of hoops, which therefore no longer provide core confinement or longitudinal bar restraint. All of the stresses are then transferred to the middle ties, which rupture. Consequently, once unbending and tie rupture has taken place, all of the longitudinal bars are no longer restrained on the edge where the unbending occurred, and the longitudinal bars are able to buckle. Where no unbending occurs (BE4), only the middle longitudinal bars are unrestrained (due to tie rupture), so only those bars undergo buckling and the corner bars remain effectively restrained. Because the corner bars do not buckle, the specimen was able to stay relatively vertical with very little out of plane movement. Reduction in ductility due to hoop unbending Hoop unbending occurred in all of the specimens except for BE4. As seen in Figure 13, BE4 had the highest ductility. All other specimens, which underwent hoop unbending, failed at equally low strains. This suggests that hoop unbending has an effect on the ductility of specimens. This coincides with the findings of Welt (2015), who also observed that unbending of cross ties resulted in poorer ductility than where ties didn’t unbend. Increase in ductility with uni-ties BE2 and BE4 were identically detailed with 10mm cover and 40mm transverse reinforcement spacing, however BE2 had traditional hoops and ties and BE4 had uni-ties. The two types of reinforcing arrangements are shown in Table 1. In comparing the performance of the two specimens (Figure 13), BE2 underwent hoop unbending and also displayed less ductility than BE4. Seeing as BE4 saw no hoop unbending and had a higher ductility, this would suggest that traditional hoops and ties are more susceptible to hoop unbending than uni-ties. This action, as discussed earlier, decreases the ductility. Reduction in ductility with more cover concrete BE3 and BE4 were compared for the effects of concrete cover thickness on ductility. As shown in Table 1, these two specimens are identically detailed with uni-ties and 40mm transverse hoop spacing, however BE3 has 20mm cover concrete and BE4 has 10mm cover. Considering Figure 14, strains for both of the specimens were concentrated at the location of failure. However, when looking at the spread of strain, BE4 has a much greater plastic strain spread into adjacent bays than BE3. This suggests that smaller cover concrete allows for a more even ductility spread in the specimen, and hence larger ductility overall.

With smaller cover, the confined core area of the specimen will be greater than that with larger concrete cover. Therefore, once the cover concrete begins to spall in both specimens, the core with the larger area will be able to withstand a greater force (note the stress f’cc will be the same for both specimens, but the increased area means the force applied will be greater before the confined concrete core strength is reached). Therefore, at a uniform loading rate, the specimen is able to withstand this greater force over this larger area, so more spalling will be able to occur before failure. Hence BE4 could fail in a more ductile manner. Transverse reinforcing spacing

BE4, BE5 and BE6 were identically detailed, with 10mm cover and uni-ties. These specimens were used to compare the effects of transverse reinforcement spacing on the ductility of boundary elements. The spacings were 40mm, 60mm and 80mm respectively. It was found that BE5 and BE6 both underwent hoop unbending. BE5 unbent over two levels or hoops, where BE6 unbent over only one, as shown in Figure 17. Considering that BE6 unbent over only one transverse reinforcing level, leaving an unrestrained longitudinal bar length of 160mm, and BE5 unbent over two levels to form an unreinforced longitudinal bar length of 180mm, this suggests the upper limit to cause hoop unbending is less than or equal to 160mm. The lower limit shall be 120mm, as this would be the unrestrained longitudinal bar length if only one hoop unbent in BE5. As this did not occur and BE5 needed two hoops to unbend to cause bar buckling, the unrestrained length of longitudinal reinforcement must be greater than 120mm. The absence of hoop unbending in BE4 suggests that there is a critical hoop spacing at which the failure mechanism changes from tie rupture to hoop unbending. Seeing as BE4 (with a spacing of 40mm) had no hoops unbend and BE5 (with a spacing of 60mm) did have hoops unbend, this suggest that the critical hoop spacing lies between 40mm and 60mm. This critical hoop spacing, within which the failure mechanism changes, can be represented by a critical longitudinal bar slenderness ratio s/db, of between 2.5 and 3.75. It should be noted here that NZS3101:2006 specifies a maximum s/db, ratio of 6. CONCLUSIONS From the testing carried out, the following conclusions were drawn:

• Hoop unbending can result in a change in the failure shape from a radial expansion to an out of plane failure shape, in addition to a decrease of ductility in a boundary element.

• Uni-ties provide more ductility than traditional hoops and ties, as the uni-ties are less susceptible to hoop unbending.

• With smaller concrete cover, the ductility increases, as more propagation of damage occurs over the specimen height.

• Ductility increases with smaller hoop spacing, as tighter spacing provides better confinement. The critical s/db ratio has been identified as being between 2.5 and 3.75 to avoid hoop unbending during BE failure.

Figure 17: Tie unbending in BE4, BE5

and BE6

Further research required As discussed in this report, BE4 had the most desirable response out of all specimens tested. As the conclusions of the results relied heavily BE4’s ideal behaviour, more tests on specimens with the same detailing as BE4 would be beneficial, to quantify these results and confirm consistency for this detailing. It would also be of interest to refine the critical longitudinal bar buckling ratio discussed above, by testing specimens with 50mm spacing and observing whether failure was initiated by hoop unbending or tie rupture. The specimens tested were one-half-scale boundary elements. Therefore, full-sized specimens could be tested to check that the results are scalable. ACKNOWLEDGEMENTS

I would like to thank Dr. Chris Motter for his support and insight throughout this project. Thanks is also extended to Alex Shegay and Dr. Lucas Hogan for assisting during testing and offering advice. I would also like to thank the lab technicians: Andrew Virtue, Shane Smith, Mark Byrami and Ross Reichardt. Thanks to Holcim and Complete Reinforcing for supplying materials. Special thanks goes to my project partner, Jonathan Wood for his dedication, bad jokes and positivity during the project. REFERENCES Arteta, C. A. (2015). Seismic Response of Thin Boundary Elements of Special Concrete Shear

Walls. (Doctor of Philosophy Dissertation). University of California, Berkeley, University of California, Berkeley.

Hilson, C., Segura, C., & Wallace, J. (2014a). Experimental Study of Longitudinal Reinforcement Buckling in Reinforced Concrete Structural Wall Boundary Elements.

Mander, J., Priestley, M., & Park, R. (1988a). Oberserved stress-strain behaviour of confined concrete. Journal of Structural Engineering, 114(8), 1827-1849.

Massone, L., Polanco, P., & Herrera, P. (2014). Experimental and Analytical Response of RC Wall Boundary Elements.

Standards New Zealand. (2006). Concrete Structures Standard, Part 1 - The Design of Concrete Structures. (Wellington, New Zealand: Standards New Zealand).

Welt, T. (2015). Detailing for compression in reinforced concrete wall boundary elements: experiments, simulations, and design recommendations. University of Illinois at Urbana-Champaign.