édération Internationale du Béton Proceedings of the nd International Congress ID 8-34 Session 8 Seismic Behaviour of the Column oundation Connection of Pre-cast Industrial RC rames Dimova, S.L. Central Laboratory for Seismic Mechanics and Earthquake Engineering, Bulgarian Academy of Sciences, bl.3 Acad. G. Bonchev str., Sofia 1113, Bulgaria Currently: European Laboratory for Structural Assessment, Institute for the Protection and Security of the Citizen, Joint Research Centre of the European Commission, T.P. 48, I-1 Ispra (VA), Italy Negro, P., Pinto, A. European Laboratory for Structural Assessment, Institute for the Protection and Security of the Citizen, Joint Research Centre of the European Commission, T.P. 48, I-1 Ispra (VA), Italy INTRODUCTION The study is based on the experimental results for the seismic response of a pre-cast single storey RC industrial frame and on the cyclic tests of pre-cast columns. The frame structure was tested in the European Laboratory for Structural Assessment (ELSA) of the Joint Research Centre (JRC) of the European Commission at Ispra, in the framework of the research project Seismic behaviour of reinforced concrete industrial buildings by means of the Ecoleader programme, which was reserved to the European Consortium of Laboratories for Earthquake and Dynamic Experimental Research. The objective of the project was to provide specific experimental evidence about the seismic behaviour of pre-cast single-storey frames for industrial buildings as compared to the corresponding cast-in-situ analogous structures. The results were expected to contribute to the correct calibration of Eurocode 8 design rules. or this purpose, prototypes of pre-cast fand of cast-in-situ single-storey frames have been designed, both consisting of six columns connected by two lines of beams and an interposed slab. The pre-cast columns were tested in ELSA under a contract with Assobeton in 1995-1998. The experimental data for the industrial pre-cast frame have shown that under seismic excitations the curvatures below the top of the foundation exceed those above it. The records of the relative displacements at the bottom of the instrumented columns are used to estimate the contact forces acting on the faces of the columns connected to the pocket. A refined model of the column-foundation connection, which agrees well with its experimentally estimated dynamic behaviour is presented. It makes possible to introduce a simplified expression for the pressing force on the pocket faces, needed for the assessment of the horizontal reinforcement around the top half of the pocket. The expression gives well coinciding with the experimental data and slightly conservative results for the maximum value of the pressing force and could be easily implemented in the design practice. Keywords: pre-cast frame structure, column-foundation connection, seismic design, Eurocodes EXPERIMENTAL PROGRAMME The experimental prototype shown in igure 1a was designed according to Eurocode [1] and Eurocode 8 [] for ductility class, as described in details in the design calculations [3]. The two bays were 4 m each, the storey height was 5.5 m. The columns (cross-section 3mm X 45 mm) have been reinforced with 8 φ16 bars, stirrups φ6@5 mm in the critical zones and φ6@15 mm in the central part. The beams (crosssection 3mm X 6 mm) were reinforced with 8 φ14 bars and stirrups φ6@15 mm. The slab consisted of 15 mm deep hollow-core pre-cast elements connected by cast-in-situ longitudinal joints and tied together by a peripheral curb, cast along the deck perimeter and provided with minimum reinforcement. Neoprene pads, 6 mm thick, were inserted between the column top edges and the beams, providing for the hinged support, φ6 dowels were used for the horizontal connections. The foundations consisted of six pre-cast blocks
anchored to the floor. The columns were inserted in the pockets of the blocks and fixed with shrinkagecompensated mortar [4]. ig. 1a. View of the experimental frame The structure was designed for compressive strength of concrete f ck = 4 MPa and yielding strength of steel (type B5) f yk = 5 MPa. The measured mean value of the cylindrical compressive strength was f cm = 43. MPa. The results from the tests on longitudinal reinforcement samples (φ16) have shown yielding strength f y = 55 MPa and tensile strength f t = 657 MPa [4]. The pseudodynamic tests have been carried out in ELSA in September [4-7]. The denotations adopted for description of the experimental results and data about the disposition of the transducers are shown in igure 1b. In order to reproduce the design values of the axial loads in the columns, a total vertical load of 6 kn was applied by means of vertical jacks, as shown in igure 1a. The horizontal displacements were imposed to the structure by means of hydraulic actuators connected with spherical joints to the deck, in the vicinity of the midspan of the first bay (see igure 1a). The resulting reaction forces were measured by load cells, placed in the actuators. The structure was instrumented by relative displacements transducers placed on six levels on the North and South faces at the bottom of columns NE, NC, NW and SC (see igure 1b and igure 1c). On the basis of the measurements of the relative displacements, the mean curvatures for each segment were calculated as the estimated cross-section rotation divided by the corresponding gauge length. L 6 N Applied displacement SE SC SW NE NC NW Applied displacement R E A C T I O N W A L L 1..85.4 L 5 L 4 L 3 L L 1.376 6 X.15 foundation ig. 1b. Plan of the experimental structure ig. 1c. Instrumentation of the bottom part of the columns
The seismic excitation applied exhibits Eurocode 8 compatible design spectrum, depicted in igure. 3 Sa/PGA 1.4.8 1. 1.6 T, s ig.. Acceleration response spectrum of the experimental accelerogram The experimental programme comprised pseudodynamic tests with peak ground acceleration (PGA) of 5% g, 36% g and 7% g, as well as an additional excitation with PGA of 18% g, where g is the acceleration of gravity. During the latter test the amplitude of motion took the hydraulic actuators to the end of their stroke and the test had to be stopped. Only some information was recorded in the beginning of the test [4]. EXPERIMENTAL RESULTS The maximum mean curvatures at Levels 1, and 3 for the tests with PGA of 5% g, 36% g and 7% g are presented in Table 1. or the excitations with PGA of 36% g and 7% g, the largest curvatures were registered at Level 1, which is below the top of foundation. Only in the case of column NW, the maximum curvature at Level is larger than those at Level 1 for the excitation with PGA of 7% g. or the same column the records at Level 1 for the excitation with PGA of 36% g were unreadable. Tab. 1. Maximum curvatures measured column Level φ max, rad/m 5% g 36% g 7% g 1.696.165.757 NE.184.111.364 3.75.1373.456 1.968.6.815 NC.8.1687.6 3.194.14.43 1.595 -.591 NW.59.1.719 3.5.147.7 1.863.19.64 SC.161.131.53 3.118.134.48 The existence of significant curvatures at the level below the top of the foundation has been experimentally proven in the studies of the dynamic behaviour of a cast-in-situ industrial frame [9] as well as of pre-cast columns [1-14]. This phenomenon is referred to as tensile strain penetration [15] since the steel tensile strains continue, due to finite bond stress, for some depth inside the footing. owever, in the present study, the curvatures below the top of the foundation exceed those above it and this phenomenon cannot be explained only by the tensile strain penetration. 3
.1 D Proceedings of the nd Congress Session 8 EXISTING MODELS O TE COLUMN- OUNDATION CONNECTION The overview of the studies on the pre-cast structures [16-19] shows that, as mentioned in [16] there is a lack of analytical or experimental data on the real behaviour of pocketed connections, but this is most probably due to an almost total absence of failures. The only research on this topic has been limited to considering the prevention of concrete splitting in the sides of the pocket [19]. The diagonal-tension shear across the corner of the pocket (for which links are provided around the top half of the pocket) and crushing of the in-situ grout in the annulus are considered in [16] as possible failure modes. The model presented in igure 3 was developed in [16] for the case of shear force present in the column. According to it, the top.1 D part of the pocket is ignored within the cover zone. An ultimate horizontal compressive stress of.4 f cu b is taken in the in-situ grout across the breadth of the column, where f cu is the strength of the infill and b is the width of the column. The shear force is considered as transmitted in the upper part of the pocket along the length L defined as: L.4 f ' b = (1) cu where L is measured from a point at.1 D from the top of the pocket, as shown in igure 3. e N A µ L 3 L D L 3 µ h ig. 3. Elliot s model of the column-foundation connection The moment is transferred from the column to the foundation by a set of pressing contact forces and vertical friction forces µ, where µ =.7 is accepted in [16]. The pressing forces and the respective friction forces µ are considered to act only over the lengths L 3, as shown in igure 3. The pressing forces are defined as:.4 f ' L b = () cu 3 where L 3 is calculated from the moments equilibrium about point A (see igure 3): Ne +.1D +.5L ) =.4µ f ' bhl +.4 f ' bl (.9D L ) (3) ( cu 3 cu 3 L3 In this way the force is calculated and the horizontal force + to be used for computing the horizontal reinforcement needed around the top half of the pocket can be estimated (in the case of tapped pockets also the wedge force is included [16]). 4
It should be noted, that during the experimental study neither cracking across the corners of the pocket, nor crushing of the in-situ grout in the annulus was observed. Also, the transmission of the shear and contact forces in the Elliot s model is assumed to occur in two separate zones in the upper part of the pocket. The associated presence of friction (and the respective possibility for sliding of the surfaces) only in the lower zones with length L 3 does not explain the large relative displacements in the column at Level 1. ence, the above model in its current state cannot explain the considerable curvatures registered below the top of the foundation and needs further development. ESTIMATION O TE CONTACT ORCES The experimental data show that in all columns the relative displacements at Level 1 become commensurate or larger than those at Level during the test with PGA of 36% g. This fact is illustrated in igure 4 a, b, where the relative displacements of column SC, side NE are shown for the tests with PGA of 5% g and 36% g, respectively..16.1 column SC, side NE. 5%g Level Level 1.8.6 column SC, side NE, 36% g Level Level 1.8.4.4. -.4 -.8 4 8 1 16 -. 1 3 ig. 4a. Relative displacements of column SC (side NE) for PGA of 5% g ig. 4b. Relative displacements of column SC (side NE) for PGA of 36% g The records of the relative displacements at Level 1 during the test with PGA of 36% g are used to estimate the contact forces c1 and c, acting on the faces of the instrumented columns connected with the grout. They are calculated as the external forces, which have to be applied on the two sides of the column crosssection together with the acting moment M, the shear force and the axial force N, in order to obtain the experimentally estimated mean strains (see igure 5). M N ε ε 1 c c1.3.45 ig. 5. Estimation of the contact forces 5
The shear force and the moment in the middle of Level 1 are estimated by use of the reaction forces measured in the actuators. The axial force N was obtained using the self-weight of the structure and the records of the vertical load applied to the structure. The experimental characteristics of the concrete and reinforcing steel were used. In igures 6 a, b the contact forces calculated from the records for PGA of 36% g in the South faces of columns NE and SC are shown, respectively. As it can be seen from these figures, there are considerable contact forces in the case of positive (tension) strains. or small negative strains, the surfaces of the column and the grout are coupled by considerable contact forces, which are linearly proportional to the strain. They can be represented as: = k( ε ε ) for ε ' ε (4) c where k = 865 kn is a coefficient of proportionality, which, according to the numerical results, does not depend on the axial force in the column; ε is the strain corresponding to the onset of sliding in compression. In the considered case ε = -.64 and ε = -.31 for the central columns. or the edge columns ε = -.67 and ε = -.6. 6 4 4 col. NE, side S, 36%g c1 (West face) c (East face) col. SC, side S, 36%g c1 (West face) c (East face) Contact force, kn Contact force, kn - - -4-4 -...4.6 mean strain -...4.6 mean strain ig. 6a. Contact forces in the South face of column NE ig. 6b. Contact forces in the South face of column SC The sliding in the column-grout interface starts for ε > or ε < ε ', as schematically illustrated in igure 7a. The contact forces delimiting the onset of sliding at compression cc and tension ct are estimated as cc = 8 kn and ct = -7 kn for the central columns and cc = 35 kn and ct = -3 kn for the central columns. The contact forces at tension when the shear force and the moment do not affect pressing force to this interface (see igure 3) have magnitude commensurate with that in compression. This phenomenon could be explained by the flexural deformation of the column surface, which applies pressure to the plane surface of the grout in their interaction. Despite the identical slope of the sticking branch of the contact force obtained for the edge columns (design axial force N = 9 kn) and central columns (design axial force N = 18 kn), the values of cc, ct and ε differ for the edge and central columns. The edge and central columns have identical cross-sections and reinforcement, the horizontal forces (and the respective bending moments) are also equal and the only different parameter is the axial force N. ence, the dissimilarity in the characteristics of the contact forces can be attributed to the magnitude of the axial force and to the resulting initial deformation ε i. The larger the initial deformation ε i, the larger ε and the values of the contact force are shifted on the left for ε <, as depicted schematically in igure 7b. The shifting of the sticking branch of the contact force implies larger value of the contact force ct at the beginning of sliding in tension. If c < ct when M = M y, the tension surface of the column cannot slide and no substantial curvature can be achieved below the top of the foundation. In this case the plastic hinge is developed only above the foundation. 6
c cc c cc1 cc smaller N bigger N ε ε 1 ε ε ε ε ε ε 1 ct1 ct ct ig. 7a. Scheme of the contact force strain relationship ig. 7b. Influence of the axial force on the contact force parameters The above statement is supported by the results of a set of cyclic tests of pre-cast columns [1-14]. None of the tested columns (all with cross-section 3 X 3 mm) developed below the top of the foundation curvature larger than the one above the foundation. Their initial compressive deformations ε i calculated from the experimental characteristics of the concrete were between.46 and.43. They were larger than those of the columns of the considered structure (ε i =.4 for the central columns and ε i =.11 for the edge columns). In igure 7c the ratio of the maximum curvature attained at Level 1 and Level is presented as a function of the initial strains, as reported in [1-13] and for the considered pre-cast structure. As it can be seen from igure 7 c, the larger the initial strain, the smaller the curvature below the top of the foundation. The slight increase of the curvatures for the largest ε i considered could be explained by the possibility for larger excursions in the sticking branch in case of larger ε i...5 Assobeton tests pre-cast structure ϕmaxlevel1/ϕmaxlevel 1.5 1.5 5E-5.1.15..5.3 ε ι ig. 7c. The ratio of the maximum curvature attained at Level 1 and Level The general trend of the data presented in igure 7c shows that for ε i <.43, considerable curvatures could be developed below the top of the foundation. It should be noted that to this strain corresponds the 7
.1 D Proceedings of the nd Congress Session 8 initial design strain of.54 (considering the experimentally obtained safety factor of 1.5 for the initial strains) and that present study concerns pre-cast columns with smooth (i.e. without especially roughened [16]) surface. The shape of the histeresis loops shown in igures 6 a, b proves a substantial decrease of the friction coefficient with the growth of the strains in both tension and compression sides. Also, the rate of attenuation of the friction coefficient on the tension side is much lower than those on the compression side. ence, the magnitude of the contact force corresponding to the amplitude values of the strains (and respectively the horizontal force applied) is much smaller than that in the beginning of the sliding. REINED MODEL O TE COLUMN-OUNDATION CONNECTION Based on the above-presented experimental results, a model of the column- foundation connection is considered (igure 8 a). It is assumed that the shear force is transmitted gradually along the upper half of the pocket together with the pressing force, thus making it possible the relative movement of the sliding surfaces. The pressing forces are induced by the rigid-body-like movement of the bottom of the column under the bending moment M. The contact forces in the column-grout interface act only in the zone where the column exhibits bending deformations, i.e. in the upper half of the foundation pocket. The experimental studies in [1], in which the columns were instrumented along the whole depth D of foundation pocket, have shown that bending deformations appear only in the upper third of D. M N D c c1.6 D.45 D.45 D B R h ig. 8a. Enhanced model of the column- foundation connection The moments equilibrium about the centre line of the column at point B (see igure 8 a) gives the following expression for the pressing force : M +.5D.5h( c1 c ) = (5).6D In igure 8 b, c the pressing forces obtained for column SC to the East and West side are shown. As seen from the results presented, the largest values of the pressing force correspond to the largest value of the shear force in the corresponding direction. The obtained pressing forces for the cases of small shear with 8
opposite sign could be attributed to the moments created by the contact forces of the uploading hysteresis branches, which oppose the moment induced by the horizontal seismic loading. 5 Col. SC, 36%g pressing force to the E side contact forces considered design fomula 5 Col. SC, 36%g pressing force to the W side contact forces considered design fomula 4 4 Pressing force, kn 3 Pressing force, kn 3 1 1-6 -4 - Shear force, kn - 4 6 Shear force, kn ig. 8b. Pressing force to the East side of column SC ig. 8c. Pressing force to the West side of column SC Equation (5) involves the contact forces in the column-grout interface and in this form is inappropriate for use in the design practice. The conclusion that the magnitude of the contact force corresponding to the amplitude values of the horizontal force applied is much smaller than that in the beginning of the sliding (presented in section Estimation of the contact forces ), was used to simplify Equation (5) in the form: M +.5D = (6).6D In igures 8 b, c the results obtained by use of Equations (5) and (6) are compared. The simplified formula gives well coinciding and slightly conservative results for the maximum value of the shear force in the respective direction, which is the design case. In the case when no sliding appears in the column-grout interface, Equation (6) also gives conservative results, since the second term in Equation (5) is neglected. This statement is illustrated in igures 8 b, c, where the results from Equation(6) provide an upper bound of the results from Equation (5) for smaller values of the shear force. Using Equation (6) a check of the crushing of the grout during the test with PGA of 7% g has to be performed, when the maximum shear force induced in the columns reached 53.68 kn. A strength of the infill f cu = 4.1 3 kn/m is considered according to [16]. The maximum compressive stress p max according to the model illustrated in igure 8a is expressed as: p max ( + ) = (7).45Db where f cu is the strength of the infill and b is the width of the column. An ultimate horizontal compressive stress of.4 f cu b is taken in the in-situ grout across the breadth of the column. The obtained p max = 1585 kn is smaller than f cu and validates the lack of crushing of the grout after the test with PGA of 7% g. CONCLUSIONS 1. The pre-cast columns with smooth (i.e. without especially roughened) surface can develop considerable curvatures below the top of the foundation. Based on the considered experimental results, one may set the upper boundary of the initial compression design strains when considerable curvatures below the top of the foundation occur, as.54. 9
. The presented refined model of the column-foundation connection agrees well with the experimentally assessed behaviour. This gave the possibility to propose a simplified expression for the pressing force on the column faces. The expression provides results which are well coinciding with the experimental data and slightly conservative and could be easily implemented in the design practice. ACKNOWLEDGEMENTS The test results mentioned in the present paper were obtained as a part of the research project Seismic Behaviour of Reinforced Concrete Industrial Building, funded in the V ramework Programme under the contract ECOLEADER. S. Dimova is a JRC Seconded National Expert from CLSMEE - BAS, Letter No B1- R/ERO/lca/D(5)5351. The support of the JRC/European Commission is acknowledged. REERENCES 1. EN199-1-1. Eurocode : Design of concrete structures. Part 1-1: General rules and rules for buildings. Commission of the European Communities, European Committee for Standardization, Draft January 1993.. ENV 1998-1-1. Eurocode 8: Design of structures for earthquake resistance. Part 1: General rules, seismic actions and rules for buildings. Commission of the European Communities, European Committee for Standardization. pr Draft 3, May 1. 3. errara L. Design calculations. ECOLEADER research project: Seismic behaviour of reinforced concrete industrial buildings. Politecnico di Milano;. 4. errara L, Negro P. Seismic behaviour of reinforced concrete structures: test on the pre-cast prototype. Research Report EUR 196 EN, European Commission, Joint research Centre, IPSC, 4. 5. Molina J, Verzeletti G, Magonette G, Buchet Ph, Géradin M. Bi-directional pseudodynamic test of a fullsize three-storey building. Earthquake Engineering & Structural Dynamics 1999; 8:1541-1566. 6. Molina J, Verzeletti G, Magonette G, Buchet Ph, Renda V, Géradin M, Parducci A, Mezzi M, Pacchiarotti A, ederici L, Mascelloni S. Pseudodynamic tests on rubber base isolators with numerical substructuring of the superstructure and strain-rate effect compensation. Earthquake Engineering & Structural Dynamics ; 31:1563-158. 7. Molina J, Magonette G, Pegon P. Assessment of systematic experimental errors in pseudodynamic tests. In: Proc. 1th European Conference on Earthquake Engineering,, Elsevier Science, Paper 55. 8. Biondini, errara L, Negro P, Toniolo G. Results of pseudodynamic test on the prototype of a pre-cast RC frame. In: Proc. 14 th C.T.E. Congress, Mantova, Italy, November. 9. Dimova S, Negro P. Influence of the quality of construction on the seismic vulnerability of structures. Research Report EUR 19 EN, European Commission, Joint research Centre, IPSC, 4. 1. Saisi A, Negro P. Programma di prove ASSOBETON. Relazione n.1. Special Publication No I.95.45, European Commission, Joint research Centre, ELSA, 1995. 11. Saisi A, Verzeletti G, Negro P. Programma di prove ASSOBETON. Relazione n.. Special Publication No I.96.5, European Commission, Joint research Centre, ELSA, 1996. 1 Saisi A, Verzeletti G, Negro P. Programma di prove ASSOBETON. Relazione n.. Special Publication No I.96.5, European Commission, Joint research Centre, ELSA, 1996. 13. Saisi A, Verzeletti G, Negro P. Programma di prove ASSOBETON. Relazione n.3. Special Publication No I.98.4, European Commission, Joint research Centre, ELSA, 1998. 14. Saisi A, Verzeletti G, Negro P. Programma di prove ASSOBETON. Relazione n.4. Special Publication No I.98.41, European Commission, Joint research Centre, ELSA, 1998. 15. Paulay T, Priestley MJN. Seismic design of reinforced concrete and masonry buildings. John Wiley & Sons, Inc. New York, 199. 16. Elliot KS. Multi-storey pre-cast concrete framed structures. Blackwell Science Ltd, Oxford, 1996 17. Elliot KS. Research and development in pre-cast concrete framed structures. Prog. Struct. Engng. Mater, :45-48,. 18. Bruggeling AS, uyghe G. Prefabrication with concrete. Balkema, Rotterdam, 1991. 19. Somerville G. Tests on column-column joints for the Ministry of public building and works. Cement & Concrete Association, Wexham Springs, Note DW/3, July 1967. 1