W2A-035 Concrete slabs-0.txt
Post-Tensioned Concrete Flat Slabs at Edge Columns

by Adrian E. Long and David J. Cleland

An experimental study of the behavior of post-tensioned unbonded concrete flat slabs is presented. This describes the testing of five model slabs that closely simulated an exterior panel in a real structure but at a reduced scale. The paper discusses the behavior at three stages: during transfer, service conditions, and at ultimate load. The ultimate capacities are compared with current design provisions, and modifications are recommended.

Keywords: columns (supports); concretes; concrete slabs; connections; flat concrete slabs; models; post-tensioning; prestressed concrete; tests.

Flat slabs offer an efficient and economical method of construction for commercial and industrial buildings. Post-tensioning in place of reinforced concrete leads to further advantages in economy and versatility. The enhanced flexural characteristics of post-tensioning provide the potential for longer spans without increasing the slab depth. However, flexural failure is not the only limit state to be considered. Punching at the slab-column connections in flat slab structures is frequently more critical, and in post-tensioned slabs this is even more likely to be the case, since current building codes allow only a modest increase in the punching shear strength due to post-tensioning compared to conventional reinforcement.

ACI-ASCE Committee 4231 has recommended a simple equation for shear strength that incorporates a term for the level of average prestress. This has been based on experimental results of a small number of isolated interior column models, where a uniform distribution of shear stress around a critical perimeter has been assumed. For edge columns, the same equation for shear strength may be used with the critical shear stress obtained from a consideration of the moment transfer effect and the nonsymmetrical shear perimeter.

RESEARCH SIGNIFICANCE

Little experimental work has been carried out to investigate the validity of applying the ACI-ASCE Committee 423 approach to edge column locations in post-tensioned slabs. Tests on multipanel models at the University of Texas included some edge column failures, and a series of four tests on single isolated edge column models has been reported recently by Foutch, Gamble, and Sunidja. This paper shows that the code provisions for exterior columns are very conservative, particularly for cases where the tendons perpendicular to the edge are banded.

OBJECTIVES

An experimental investigation was initiated to study the behavior of edge panels in post-tensioned unbonded slabs with the aid of models representing the critical edge column-slab connection. Considerable importance was placed on simulating the prototype structure in every detail. The variables considered in the test series were: a) the level of average prestress, b) the distribution of prestressing within the panel, and c) the tendon profile.

TEST SPECIMENS

The test series was based on a prototype structure that is shown in Fig. 1. The 8.25-in. (210-mm) deep slab was designed on the basis of a 30.0-ft (9.14-m) span, a superimposed dead load of 15 lb/ft (0.72 kN/m), and a live load of 50 lb/ft (2.40 kN/m). To achieve full dead load balance in the interior panels, a level of average prestress of 260 psi (1.80 N/mm) was required with the tendon drape shown in Fig. 1. However, to allow for the smaller drape of the tendons in the edge panel, prestress at this location was increased to 340 psi (2.35 N/mm). Under service loads, the maximum tensile stress in the concrete would be 167 psi (l.15 N/mm) or 2.4 /f c and the maximum compressive stress 796 psi (5.48 N/mm).

Although service load tensile stresses were very small, a mat of bonded reinforcement was placed directly over the columns in the top of the slab in accordance with normal practice. Four No. 8 bars were specified in each direction.

Model configuration

Approximately 1/4-scale models of the type used by Neth were chosen (Fig. 2). The advantages of this structural model over other isolated slab-column models are as follows: a. The slab width is equal to one bay and therefore directly comparable to the section used in equivalent frame analyses b. The slab deformations are similar, if not the same, as those in the prototype. At failure, the cone of failure intersects the slab edge, whereas only torsional cracks are in evidence in models that do not have deformation similitude. c. Redistribution of moments, as a result of cracking and yielding, can take place in a manner similar to the prototype. d. The stress increase in the prestressing tendons is similar to a real structure, since stress equalization over a full panel length can take place. e. The model includes both column and middle strips, which enables the distribution of prestressed and ordinary bonded reinforcement to be studied.

The test program consisted of five models that were identical in dimension (Fig. 2). The primary variables were the average level of prestress in the slab and the use of distributed or banded tendons. In all specimens, except one, the tendon layout and prestress level were designed to approximately balance the prototype dead load of 115 lb/ft (5.60 kN/m).

The 6.4-in. (162.5-mm) square columns were supported horizontally at approximately 24 in. (600 mm) above and below the slab. Details of the slab reinforcement are presented in Fig. 3 and 4 and summarized in Table 1. Prestressing wires of 0.2-in. (5-mm) diameter coated in a water-repellant grease were used to model the 7-wire strand of the prototype. After stressing, each wire had an effective force of 4 kips (17.8 kN) in Specimen E1 and 4.50 kips (20.0 kN) in all other specimens (the number of wires being altered to compensate). To allow for the higher average level of prestress in Specimen E3, the drape of the prestressing wires was reduced relative to E2 so that the overall load-balancing effect was the same. The bonded reinforcement, No. 2 bars (6-mm diameter) over the column, was generally in compliance with Sections 18.9.3.3 and 18.9.4 of Reference 2, except that the area and number of bars perpendicular to the free edge were increased by 50 percent in Model E1, which had distributed prestress, and the length was increased by about 50 percent in Model E5.

Materials

The ultimate strength of the 0.2-in. (5-mm) diameter prestressing wires was 230 ksi (1580 N/mm). The bonded reinforcement was ribbed high-yield steel with a yield point at 75.0 ksi (517 N/mm).

The concrete was made from a 1/4-in. (6-min) maximum-size crushed basalt, a coarse grit, sand, and Type 1 portland cement. Trial mixes were made to arrive at a suitable concrete with a realistic ratio of tensile to compressive strength (Table 2).

INSTRUMENTATION AND TEST SETUP

A self-stressing test frame also capable of providing restraint to the top and bottom of the columns was used. The vertical load at the bottom of the lower column and reaction required to maintain lateral stability at the top of the upper column were transferred through roller bearings. Consequently, the horizontal reaction in the direction of moment transfer at the end of each column, and hence the bending moment, could be measured using load cells.

To model a uniformly distributed load on the slab, loads were applied at 16 points on the surface. A system of orthogonal distribution beams was used to load each half-panel of the specimen, and each was activated by a hydraulic ram reacting against the main loading beams of the test frame.

To achieve the correct boundary conditions along the continuous edges, a system of restraints was used. The slab edge was cast with discontinuities so that this restraining system prevented curvature perpendicular to the slab edge only, not parallel to the slab edge.

Force in eight typical prestressing wires was monitored using both 11.2 and 22.5-kips (50 and 100-kN) capacity load cells. Initially, these load cells were used to assess the friction loss and loss due to wedge set in each wire. This was generally around 8 percent and subsequently the jacking force was set to allow for this.

Displacement transducers were located beneath the slab to measure displacement at significant locations and five inclinometers were attached to the soffit of the slab to measure rotation along the free edge. These were not sufficiently sensitive to measure column rotation, so a vertical arm was attached to the column at slab level and the horizontal displacements at its ends were measured.

TEST PROCEDURE

A small number of prestressing wires was stressed for handling purposes before transferring each specimen to the test frame. Once in place the edge restraint system and loading beams were fixed in place while the prestress was increased to about 1/8th of the full value (the weight of the equipment and the self-weight of the slab equalled about 1/8th of the prototype dead load). From this datum, load was applied in three increments to a value equivalent to the full prototype dead load, while after each increment a commensurate increase in the prestressing was applied. Slab deflections and column moments were monitored at each increment.

The vertical live load was applied in increments up to service load and then reduced to zero (i.e., dead load only) several times to observe the serviceability behavior of each specimen and to insure that all aspects of the test setup were functioning satisfactorily. The ultimate load test was then carried out by applying the load in increments and recording all the instrumentation at each increment. In all specimens, except E1, the ends of the columns were readjusted to their original position before readings were taken to prevent error owing to movement of the column supports caused by strain of the load cells.

At several stages throughout the test the slabs were examined for cracks that were then marked and photographed.

RESULTS

Application of the design live load did not cause cracking in any of the slabs, and in each case they behaved elastically, returning to almost exactly the same position on unloading. In the ultimate load test, cracking was first observed at corners of the columns and then across the inner face, at a value of live load between 50 and 100 percent above the design value. With a further increase in load, of about 60 percent of design load (30 lb/ft>), torsion cracks had extended from the initial crack to the free edge, making an angle of about 45 deg with the free edge.

The cracking in the column region progressed in the form of radial cracks spreading from the column corners and inner face. Two cracks, generally along the lines of reinforcing bars, extended gradually to meet those originating at the opposite column (Fig. 5). Prior to failure, transverse cracks occurred on the underside of the slab midway between the columns. In Specimen E1, which had no bonded reinforcement in this region, the crack opened considerably and instantaneous deflections occurred. Although not required at service load, bonded reinforcement was incorporated at midspan in subsequent specimens, and these behaved much more satisfactorily in this respect. Six No. 2 bars were provided perpendicular to the direction of cracking. Cracks usually also occurred on the slab underside, along the line of the edge restraint, indicating significant sagging curvature at this location and consequently the effectiveness of the restraint.

In all cases, the specimens failed by one of the columns punching through the slab. This took place suddenly and was accompanied by a loud noise and particles of concrete being projected from the failure zone. The extent of the failure zone in the direction of moment transfer was about 1/6th of the span, about coincident with the ends of the bonded reinforcement. In Specimen E5, longer reinforcing bars were used, and this resulted in a more compact failure zone and one that is likely to be able to carry, by dowel action, a greater reserve load after failure.

An interesting characteristic of the failure is that it allows anchorages in the region to pull in, thereby losing most, if not all, of their force. This was particularly evident in Specimens E2, E3, and E5, all of which had banded tendons and medium to high prestress levels. The loss of prestress may, and in the case of Specimens E3 and E5 did, initiate failure at the second column.

Deflection

In each test the slab deflections increased more rapidly, due to cracking, as load increased. Specimens E1 and E4 generally exhibited the greatest deflection for any given load, since cracking tended to occur earlier and to widen more rapidly in these two tests.

The deflection profiles for Specimen E5 are shown in Fig. 6. These are typical of those obtained for other tests, but E5 was instrumented more densely. Although at the initial stages the deflection profile along the line joining the columns infers almost zero slope at both the column and midspan, the situation at higher loads is different in that there is little evidence of a contraflexure point. This is due to flexural and torsional cracking in die column region, which permits large rotation of the slab relative to the column.

The deflection profiles are all for load levels below that at which midspan cracking occurred. This is confirmed by zero slope at midspan for the deflection profiles along AB and the effectiveness of the edge restraint by the profiles along AC.

ACI member Adrian E. Long is Professor of Civil Engineering and Director of the School of the Built Environment at the Queen's University of Belfast, Northern Ireland. He is a member of joint ACI-ASCE Committee 445, Shear and Torsion. His research interests are shear of reinforced concrete and methods of in situ testing for concrete strength and durability.

AC1 member David J.Cleland is a senior lecturer in Civil Engineering at the Queen's University of Belfast, Northern Ireland. His research interests include shear in concrete stabs, concrete durability, and concrete repairs.
