EXPERIMENTAL RESEARCH OF R.C. ELEMENTS WITH SUBSTANDARD DETAILS

Transcription

1 First European Conference on Earthquake Engineering and Seismology (a joint event of the 13 th ECEE & 3 th General Assembly of the ESC) Geneva, Switzerland, 3-8 September 26 Paper Number: 819 EXPERIMENTAL RESEARCH OF R.C. ELEMENTS WITH SUBSTANDARD DETAILS Despina SYNTZIRMA 1, Georgia THERMOU 2, Stavroula PANTAZOPOULOU 3 and George HALKITIS 4 SUMMARY The inventory of structures that could be classified as old or substandard construction with reference to modern seismic design standards is vast throughout the world. For this reason development of methods for seismic assessment of such structures is a great and pressing priority for the earthquake engineering community. In the present work, an experimental research has been conducted in order to identify the interacting strength mechanisms that may develop leading individual structural components to failure, including the influence of load history on the sequence of failure. A series of component tests that comprise sixteen (16) specimens modelling reinforced concrete columns with substandard details representative of former construction practices have been tested. Columns are cantilevers with a square cross section and lap splices in the critical region. Specimens are tested in single curvature under constant axial load and cyclic lateral load reversals simulating earthquake effects. Detailing of the specimens was done so as to develop closely interacting modes of failure (shear failure after flexural yielding, lap splice failure accompanied by shear or flexural failure, etc.). This was done because based on previous research it was shown that apriori identification of the prevailing mode of failure and the associated deformation capacity is the most critical benchmark test of analytical assessment procedures of substandard reinforced concrete elements [Syntzirma and Pantazopoulou 26]. From the experimental evidence it becomes apparent that the load history played a determining role both in terms of available deformation capacity and the mode of failure that eventually prevails. The various response mechanisms such as flexure, shear, lap splice and bar buckling undergo different levels of strength degradation with increasing displacement demand and number of loading cycles. 1. DESCRIPTION OF THE EXPERIMENTAL PROGRAM The intent of the experimental program presented in this paper was to evaluate the mode of failure and dependable deformation capacity of reinforced concrete flexural frame members designed according with former design codes, which by today s standards would be considered as low ductility systems. In this regard, specimens examined were designed following the reinforcement requirements of DIN 145 of 1972, which formed the basis of design codes for reinforced concrete design in Southern Europe thirty years ago. The requirements representing seismic detailing for that chronological period were summarised for both European and North American practices [ACI ] in [Pantazopoulou 23]. The experimental program comprised 16 prismatic specimens, having a symmetrically reinforced cross section as would most likely occur in columns. The typical specimen represented a building column extending from column mid height between successive floors to the beam face at the beam-column joint interface on an interior frame connection and they were constructed at a scale of 1:2. On the basis of a survey of typical member dimensions in common residential 1 Civil Engineer, MSc, PhD Candidate, Department of Civil Engineering, Demokritus University of Thrace, Vas. Sofias 12, Xanthi 671, Greece, dsyntz@civil.duth.gr 2 Civil Engineer, MSc, PhD Candidate, Department of Civil Engineering, Demokritus University of Thrace, Vas. Sofias 12, Xanthi 671, Greece, gthermou@civil.duth.gr 3 Professor, Department of Civil Engineering, Demokritus University of Thrace, Vas. Sofias 12, Xanthi 671, Greece, tel/fax: , pantaz@civil.duth.gr 4 Civil Engineer, Department of Civil Engineering, Demokritus University of Thrace, Vas. Sofias 12, Xanthi 671, Greece 1

2 Φ12 9 Φ12 s= 5, 7, 11 or 14 (mm) direction of loading spliced bars 2 Φ6 36db or 25db 2 4Φ12 Φ6 6 Figure 1: Typical column test specimen and cross section buildings of the period considered, a single column geometry was selected that contained a variety of reinforcement details. Figure 1 depicts typical specimen details. The shear span L s was 9mm and the cross sectional dimension was 2mm square. As the focus of the project was to study only the behaviour of the columns, the foundation was made relatively stiff and strong. The longitudinal reinforcement consisted of either 8Φ12, or 4Φ12, S5 bars placed uniformly around the perimeter of the cross section, with a clear cover of approximately 2mm. Based on this parameter, specimens are separated in two groups, those having 8Φ12 referred to hereon as group I, and those reinforced with 4Φ12 referred to as group II. The number of longitudinal bars determines the shear strength demand associated with the flexural capacity, as compared with the nominal shear capacity. Thus, considering that these are specimens representative of older detailing practices (i.e. with poor transverse reinforcement), specimens in group I were made by design to be more prone to shear failure than those of group II [Syntzirma and Pantazopoulou 26a]. In contrast, specimens of group II had by a design a substantially lower shear demand than those of group II, the intent of this part of the program being to encourage the occurrence of other combinations of mixed mode failures excluding shear. Particular objective was the identification of the effects of load history on the occurrence of compression bar buckling in this group, as a symptom of premature termination of flexural response. The configuration of the transverse reinforcement is also shown in Fig.1; it was composed of smooth steel of nominal diameter Φ6, S22. Five specimens had lap splices just above the top surface of the foundation block, as is typical of building construction in southern Europe. The test specimen details are summarized in Table1. Table 1: Test Specimen Details Specimen ID Longitudinal Transverse Axial Load Lap Splice Load Reinforcement Reinforcement (P/A g f c ) l s History Spec. #1a 8 Φ 12 Φ 6 / 5.8 a Spec. #1b 8 Φ 12 Φ 6 / 5.8 b Spec. #2a 8 Φ 12 Φ 6 / 7.8 a Spec. #2b 8 Φ 12 Φ 6 / 7.8 b Spec. #3a 8 Φ 12 Φ 6 / d b a Spec. #3b 8 Φ 12 Φ 6 / d b b Spec. #4aa 8 Φ 12 Φ 6 / 7 36 d b aa Spec. #4a 8 Φ 12 Φ 6 / d b a Spec. #4b 8 Φ 12 Φ 6 / 7 36 d b b Spec. #5a 8 Φ 12 Φ 6 / 11.8 a Spec. #5b 8 Φ 12 Φ 6 / 11.8 b Spec. #7a 4 Φ 12 Φ 6 / 7.8 a Spec. #7b 4 Φ 12 Φ 6 / 7.8 b 2

3 Spec. #8a 4 Φ 12 Φ 6 / 11.8 a Spec. #8b 4 Φ 12 Φ 6 / 11.8 b Spec. #9b 4 Φ 12 2Φ 6 / 14.8 b The loading assembly used to simulate the loading conditions is illustrated in Fig.2. In the testing rig each specimen was tied back to back with a column of identical external geometry including the footing stub, which was however much stronger than the test specimen. This dummy specimen, referred to hereon as support column, was reused to conduct all tests and underwent negligible damage throughout the experimental program. The assembly of the test specimen and dummy specimen formed a composite structure resembling macroscopically a simply supported beam with a stub in the center. The ends were connected through a metal box having pinned supports to a system of transverse restrainers that allowed rotation but no transverse displacement. Transverse load was applied under displacement control in the central stub i.e. on the tied footings of the two columns, using a predefined history of displacement reversals. Axial load was applied directly on the metal box at the specimen end through a pinned support; the metal box in the other end of the assembly was attached on the testing rig through a pin, thereby transferring its reaction. The axial load was applied to its full magnitude first (8% of the column concrete crushing load A g f c ) and was subsequently maintained at that level while the system was cycled in the transverse direction. Note that as the loading sequence progressed and damage accumulated, the specimen elongated due to inelastic deformation causing a sharp increase in the axial load with increasing lateral drift; for this reason axial load was corrected in every cycle. Table 2 outlines the most important design parameters for specimens of group II; data include the calculated flexural capacity at yielding and ultimate (M y and M u ), the nominal shear demand (V u =M u /L s ) and corresponding nominal shear capacity (V n ) as well as a reduced fraction of that amount (the 65% value, which is considered to represent a critical threshold; previous studies have shown that keeping the demand below that threshold precludes premature shear failure.) Terms V s and V c represent the nominal web steel and concrete contribution to the shear strength. Table 2: Estimated strength components for specimens of group II. (M y =21 kn m, M u =26 kn m) Φ6/7 Φ6/11 2Φ6/14 ρ s, tr Shear Strength terms V s (kn) V c (kn) V n =V s +V c V u (kn m) V n Buckling Strain and Associated ductility s/φ ε buckl μ ecu Objective of the test setup adopted in the experimental program was to produce constant shear along the column spans with maximum moment at the face of the footing supports simulating the distribution of seismic moments that occur within the half span of columns in actual frame structures. The reason why there was no monolithic connection between the test column and the support column was so as to achieve realistic bond conditions and anchorage of the longitudinal bars in the test column. The relative displacements were measured for both the test and the support specimen from fixed reference points using Linear Variable Displacement Transducers (LVDTs). Data were acquired and stored using a personal computer and data acquisition system. Furthermore a grid of optical targets was placed in the critical zones of the two columns and was monitored with digital cameras throughout the test as shown in Fig.3. Displacements and surface concrete strains were determined after processing of the digital images from the target movement and grid distortion. (Image processing was done using software specific to this purpose that has been developed at Demokritus University). A parameter of the investigation was the type of displacement history applied in the transverse direction. For this reason two different displacement histories were considered in the entire test program (Fig.4): in the first displacement history, identified herein as a or aa, the displacement amplitude was increased gradually, the amplitude of each cycle given as a fraction or a multiple of the estimated yield displacement of the test specimen, Δ y. Thus, three cycle increments were applied at each displacement level in both directions; as.25δ y,.5δ y,.75δ y, 1.Δ y, 1.5Δ y, 2.Δ y, 2.5Δ y and 3.Δ y. In most cases specimens failed before the 3

4 last level was attained. The second load history, identified as b, resembled near field earthquake effects with a significant increase of displacement from cycle to cycle (.5Δ y, 1.Δ y, 2.Δ y, 5.Δ y ). The load history was explored as a test parameter because, based on earlier research results, it has been shown that this parameter determines the ductility level upon which buckling of longitudinal reinforcement may occur [Syntzirma and Pantazopoulou 26b]. In the absence of well anchored, stiff and closely spaced ties, the strain ductility of compression reinforcement is associated with sideways buckling as the expected mode of failure (due to cracks remaining open in the plastic hinge region owing to cyclic shear reversal). The relationship controlling the critical spacing of ties and the reinforcement stress that may be sustained prior to buckling is given by [Pantazopoulou 23], Test Support Axial Load LVDT LVDT LVDT LVDT LVDT Applied Cyclic Moment Distribution Figure 2: Loading assembly s / Db =.785 Es f s ; Es = 2 GPa, if f s f y ; Es = Er, if f s > f y (1) where E r is the double modulus stiffness of the reinforcement. This is a weighted average between the tangent stiffness of the bar, E h, and the initial elastic modulus E, intended to account for the elastic unloading of the tension side of the buckling bar as it bends. The equivalent axial compressive strain at which reinforcement is likely to buckle, ε buckl, is calculated from Eq. (1) and clearly depends upon the strain hardening characteristics of compression steel. The above relationship breaks down if E h =. For the bar to survive through the yield plateau region and to enter strain hardening, the adjacent core must be effectively confined so as to enable load redistribution from the bar to the concrete once the bar stiffness decreases as a consequence of buckling. The axial compressive strain that corresponds to the deformation capacity of the confined core is estimated from the normalized effective confining pressure in the direction of lateral sway, k e ρ s,st f yst /f c using pertinent confinement models [Syntzirma and Pantazopoulou 26b, Pantazopoulou 23]. Thus, the transverse reinforcement uniquely determines the critical compression strain that the core may sustain without loss of integrity; for a given s/d b ratio, the critical compression strain in moment curvature analysis should not exceed the limit ε crit =min{ε buckl,ε cu }, where, ε buckl is the critical buckling strain obtained from Eq. (1) and ε cu is the strain in the compression reinforcement when the maximum core strain reaches the limit given by the confinement model (Eq. 2). o cu ε = ε cu 24.6k eρs, (1 + ' f c st f yst ) ;.3 ε o cu.4 (2) 4

5 This value is used to evaluate M buckl from monotonic flexural strength calculations. Using the stress strain properties of the longitudinal reinforcement, the critical buckling strain was estimated for the present study in the range of (Table 2). The associated drift at compression bar buckling depends on the loading history. It has been shown that for symmetric displacement histories, the curvature ductility, μ φ, corresponding to the attainment of the critical compression strain ductility μ ε,buckl is [Pantazopoulou 23], μφ = (με,buckl + 1) ( 1 x/d) ; x: depth of compression zone (3) Similar expressions may be derived for other (non-symmetric) displacement histories. Eq. 3 is valid for the tests of the present investigation (symmetric loading). Given the relatively low magnitude of the applied axial load (x/d <.2), and the relatively sparse spacing of stirrups the estimated strain ductility for the test specimens is in the range of 2, whereas the corresponding curvature ductility upon bar buckling is estimated using Eq. (3) as 2.5. This implies that although specimens in group II are expected to survive shear failure (owing to the low demand), they are expected to fail by buckling of compression reinforcement at a low displacement ductility mm Figure 3: Position of targets placed on the specimens A target in detail Load History a Load History aa Load History b Figure 4: Lateral loading histories PRESENTATION OF TEST RESULTS In the following sections the primary experimental results are outlined for the group II (Specimens 7a,b, 8a,b, and 9a,b); the experimental evidence concerning the behaviour of specimens of group I has been described in detail in [Syntzirma and Pantazopoulou, 26a]. Observations from the failure mode characteristics are included along with the damage patterns recorded on the specimens. Yield displacement used to control the load history was estimated considering contributions from flexure, shear and pullout, in the range of 14 mm [Syntzirma and Pantazopoulou 26a, b]. Thus, the ductility level attained at the various stages of the tests corresponds to fixed multiples of this amount. Displacement capacity was identified as the displacement limit associated with a 2% drop of strength in the post-peak branch of the envelope curve, which was drawn by connecting the peak points of the experimental hysteresis loops. 2.1 Specimen #7a As was intended by the amounts of longitudinal reinforcement that were provided during design, Specimen #7a exhibited initially mostly flexural response, up until a displacement ductility magnitude of μ Δ =2.5 (first cycle at that level). Distortion in the plastic hinge zone with intense dowel action took place at that instant, as illustrated in Figure 5a and the cover started to disintegrate due to sideways buckling. When the displacement level μ Δ =2.5 was reached for the third time, the concrete cover was smashed and yet the strength of the specimen was substantial until the displacement level of μ Δ =3. was reached for the second time (Figure 5b). Buckling of the longitudinal reinforcement was the predominant mode of failure, as anticipated by the analytical model. Figure 5 also shows the hysteretic diagram of load vs applied drift for Specimen #7a. 5

6 Spec. #7a 6 P (kn) 3 θ (rad) -,2 -,1,,1, (a) (b) Figure 5: Load vs drift diagram for Specimen #7a; (a) Picture of the specimen at μ Δ =2.5, for the first time Dowel action; (b) Picture of the specimen at failure (μ Δ =3. for the second time) 2.2 Specimen #7b Specimen #7b had insignificant flexural cracks until a displacement level of μ Δ =2. was reached. During the next cycle at μ Δ =5., the first shear cracks were formed and the concrete cover started to disintegrate (Figure 6a). During the second cycle of applied displacement to ductility μ Δ =5., buckling of the longitudinal reinforcement prevailed leading to severe loss of strength (Figure 6b). 2.3 Specimen #8a Shear cracks started to form in Specimen #8a, when the displacement level of μ Δ =2.5 was reached for the first time. On the next cycle in the same displacement level the concrete cover started to disintegrate rapidly. In the end buckling of compression reinforcement occurred at a displacement level μ Δ =3. (for the first time). 2.4 Specimen #8b Specimen #8b exhibited an initially shear controlled behaviour. The first diagonal cracks were formed at a displacement level μ Δ =2., as Figure 8a indicates. Yet shear failure combined with buckling of the longitudinal reinforcement lead to severe cracking and failure at a displacement level μ Δ =5. for the first time (Figures 8b, 8c). 2.5 Specimen 9b Specimen #9b exhibited initially mostly flexural response up until a displacement level of μ Δ =2.. A few insignificant shear cracks were formed in the beginning of the next cycle, at a ductility level of μ Δ =5.. Beyond that point, bar buckling of the reinforcement prevailed as shown in Figures 9a and 9b. Failure occurred at the second cycle at displacement ductility of μ Δ =5.. 6

7 Spec. #7b 6 P (kn) 3 θ (rad) -,2 -,1,,1, (a) (b) Figure 6: Load vs drift diagram for Specimen #7b; (a) Picture of the specimen at μ Δ =5., for the first time; (b) Picture of the specimen at failure (μ Δ =5. for the second time) Spec. #8a 6 P (kn) 3 θ (rad) -,2 -,1,,1, Figure 7: Load vs drift diagram for Specimen #8a; Picture of the specimen at failure (μ Δ =3. for the first time) 3. DISCUSSION OF THE TEST RESULTS Sixteen half scale reinforced concrete columns were tested under reversed cyclic lateral displacements and axial load in the laboratory. The columns had details typical of those found in pre 197 s construction, including light transverse reinforcement and lap splices in the plastic hinge region. The axial load was.8a g f c and the load history was a parameter of the investigation. The present work has focused only on the specimens that failed due to buckling of the longitudinal reinforcement. Test observations are summarized as follows: -Specimens were not sufficiently confined as the transverse reinforcement comprised smooth Φ6 bars at a 7

8 spacing of 7mm or 11mm or 14mm. Although stirrups were sparsely spaced, shear demand was kept low by design so as to preclude the occurrence of premature shear failure, in order to observe termination of flexural response by reinforcement buckling. -Due to the wide spacing of stirrups, the unsupported length of compression rebars was significant (s/φ ratios of almost 6, 9, and 12), hence the clamping action provided by the stirrups could not support the bars under compression. Specimen 9b had double stirrups at twice the spacing of those of specimen 7b, thus the nominal shear strength in the two cases was identical; also in terms of lateral support provided to the bars the extensional stiffness of the stirrups in the two groups was identical. Both specimens reached a displacement ductility of about 5, two times before incipient failure. On the contrary, specimen 8b with an s/φ ratio of 9 but a lower extensional stiffness of the system of stirrups than either 7b or 9b failed by bar buckling at a displacement ductility of 3. -Buckling occurred in all cases tested. Axial load caused significant degradation on the overall capacity of specimens. Note however that Specimens #7a and #7b, which had transverse reinforcement at a spacing of 7mm demonstrated a more controlled state of damage involving greater areas within the plastic hinge region; in contrast specimens with wider stirrups experienced a much more brittle mode of failure. Spec. #8b 6 P (kn) 3 θ (rad) -,2 -,1,,1, (a) (b) (c) Figure 8: Load vs drift diagram for Specimen #8b; (a) Picture of the specimen at a displacement level μ Δ =2.; (b) and (c) Pictures of the specimen at failure (μ Δ =5. for the first time) 3. CONCLUSIONS In this paper results of an experimental program were presented, comprising models of reinforced concrete columns representative of old construction (sparse, inadequately anchored stirrups). Specimens were loaded under combined axial loading and reversed cyclic lateral displacements simulating seismic load. Specimens were under-designed in flexure so as to promote a flexural mode of failure despite the sparse stirrup spacing. Objective was to isolate the effect of widely spaced stirrups in promoting buckling of compression reinforcement while avoiding a mixed mode shear failure, so as to identify the limiting deformation capacity associated with the buckling mode of failure. A parameter of this and previous related studies on reinforced concrete members with brittle details was the influence of the loading history (the pattern of displacement increase from cycle to cycle), which was intended to illustrate the damaging potential of near field earthquakes (Syntzirma and 8

9 Pantazopoulou 26a]. All specimens failed in a brittle mode, invariably involving buckling of compression reinforcement, consistent with the intended design of the experimental program. From the experimental evidence it became apparent that the load history played a determining role both in terms of available deformation capacity and the mode of failure that eventually prevailed (particularly with regards to the level of displacement ductility that limited the flexurally-dominated response due to bar buckling). From the tests it is evident that the various response mechanisms of resistance, namely flexure, shear and bar buckling undergo different levels of strength degradation with increasing displacement demand and number of loading cycles. So, even if an intrinsic hierarchy of strengths is designed for at flexural yielding, it was demonstrated through the tests that this objective can be entirely reversed at higher levels of imposed ductility. Naturally, once the capacity of the weakest resistance mechanism is exceeded failure occurs regardless of the strength reserves available in all other mechanisms of resistance. The onset of localization in a specific response mechanism of a reinforced concrete element limits the dependable deformation capacity of the member. Spec. #9b 6 P (kn) 3 θ (rad) -,2 -,1,,1, (a) (b) Figure 9: Load vs drift diagram for Specimen #9b; (a) Picture of the specimen at μ Δ =5., for the first time; (b) Picture of the specimen at failure (μ Δ =5. for the second time) 5. ACKNOWLEDGEMENTS Research presented in this paper was funded by the Hellenic General Secretariat for Research and Technology, through the Project ARISTION. 6. REFERENCES ACI , (1963), ACI Standard: Building Code Requirements for Reinforced Concrete. American Concrete Institute, Detroit, Michigan, pp. 141 DIN 145, (1972) Beton und Stahlbetonbau, Bemessung und Ausführung. Deutsches Institut für Normung, Berlin 9

10 Imran I. and Pantazopoulou S. J. (1996) Experimental Study of Plain Concrete Under Triaxial Stress, ACI Material Journal, Vol. 93, No. 6, Nov.-Dec. 1996, pp Pantazopoulou S.J., (23), Chapter 4 in Seismic Assessment and Retrofit of Reinforced Concrete Buildings, fib Bulletin No. 24, Case Postale 88, CH-115, Lausane, Switzerland Pujol S., Sozen M. and Ramirez J., (26), Displacement History Effects on Drift Capacity of Reinforced Concrete Columns, ACI Structural Journal, V. 13, No. 2, March-April 26, pp Syntzirma, D.V. and Pantazopoulou, S.J. (25), Drift Capacity of R.C. Members with Substandard Details, Proceedings of the fib Symposium Keep Concrete Attractive, Paper 346, May 23-25, 25 Budapest, Hungary Syntzirma, D.V. and Pantazopoulou, S.J. (26a), Assessment of Deformability of R.C. Members with Substandard Details, Proceedings of the 2 nd International fib Congress, Paper 446, June 5-8, 26 Naples, Italy Syntzirma D. V. and Pantazopoulou S. J., (26b), Deformation Capacity of R.C. Members with Brittle Details Under Cyclic Loads, ACI Special Publication, Cyclic Shear, edited by ACI Committee 445 (Shear and Torsion), in press, American Concrete Institute, Farmington Hills, MI, USA. 1