Western Ontario University, Canada – Research Study

Novel toe driving for thin-walled piles and performance of fiberglass-reinforced polymer (FRP) pile segments

Mohammed Sakr, M. Hesham El Naggar, and Moncef Nehdi

Abstract: Despite the rapidly growing use of pile foundations, it is presently difficult to assure the integrity and uniformity of the cross-sectional area of cast-in-place piles when using normal concrete. Cavities and soil encroachments leading to soil pockets can jeopardize their load-bearing capacity. Moreover, corrosion in reinforced concrete and steel shell piles has been very costly, exceeding US$2 billion in annual repair costs in the United States alone. To address these two challenges, extensive research has been underway at the University of Western Ontario to develop novel technology for the construction of piles. Self-consolidating concrete (SCC), a material that flows under gravity and assures the integrity of piles, is cast into fiberglass-reinforced polymer (FRP) tubes that provide corrosion-resistant reinforcement. A toe driving technique was developed to install the empty FRP shells into the soil, and SCC is subsequently cast into the shells. Driving tests using this new technique were carried out on large-scale model FRP and steel pipe piles installed in dense dry sand enclosed in a pressure chamber. FRP–SCC and steel closed-end piles were also driven using conventional piling at the pile head. Static load tests were conducted on the various pile specimens under different vertical and horizontal confining pressures. The pile specimens were instrumented to investigate their dynamic behaviour under driving and their response to static compressive, uplift, and lateral loading. It is shown that the toe driving technique is very suitable for installing FRP piles in dense soils.

Results from the driving tests and static load test indicate that FRP–SCC hybrid piles are a very competitive and attractive option for the deep foundations industry.

Key words: FRP, self-consolidating concrete, piles, pile drivability, toe driving, axial load, uplift load, lateral load, large-scale modeling, shaft resistance, dense sand.

Résumé : En dépit de l’utilisation croissante des fondations de pieux, il est présentement difficile de s’assurer de l’intégrité et de l’uniformité de la section en travers des pieux coulés en place lorsqu’on utilise le béton normal. Des cavités et des empiétements du sol conduisant à des poches de sol peuvent compromettre la capacité portante de ces pieux. De plus, la corrosion des armatures du béton armé et des enveloppes d’acier des pieux a causé des coûts très importants excédant 2 milliards de dollars en coûts de réparation annuelle aux États-Unis seulement. Pour s’attaquer à ces deux défis, une recherche intense a été entreprise à la University of Western Ontario pour développer une nouvelle

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Key: Blue = notable highlights Red = important information = critical informationHighlights added by Tom Anderberg

technologie pour la construction de pieux. Un béton auto-consolidant (SCC), matériau qui s’écoule sous l’effet de la gravité et assure l’intégrité des pieux, est coulé dans des tubes de polymère renforcé de fibre de verre (FRP) qui four-nit une armature résistante à la corrosion. Une technique de fonçage par le pied a été développée pour installer dans le sol les tubes vides de FRP, et par la suite, le SCC est coulé dans ces tubes. Des essais de fonçage au moyen de cette nouvelle technique ont été réalisés sur des modèles à grande échelle de pieux avec tubes FRP et tubes d’acier mis en place dans un sable dense sec à l’intérieur d’une chambre à pression. Des pieux de FRP–SCC et d’acier à bout fermé ont aussi été foncés au moyen de fonçage conventionnel sur la tête du pieu. On a fait des essais de chargement statique sur les divers spécimens de pieux sous différentes pressions de confinement verticales et horizontales. Les spécimens de pieux ont été instrumentés pour étudier leur comportement dynamique durant le fonçage et leur réaction à des chargements statiques en compression, en soulèvement et latéraux. On montre que la technique de fonçage en pied est très adéquate pour la mise en place des pieux FRP dans les sols denses. Les résultats des essais de fonçage et de l’essai de chargement statique indiquent que les pieux hybrides FRP–SCC sont une option très attrayante et compétitive pour l’industrie de fondations profondes.

Mots clés : FRP, béton auto-consolidant, pieux, potentiel de fonçage, charge axiale, résistance au soulèvement, charge latérale, modèle à grande échelle, résistance du fût, sable dense.

[Traduit par la Rédaction]

Received 20 December 2002. Accepted 3 September 2003. Published on the NRC Research Press Web site at http://cgj.nrc.ca on 14 April 2004.

M. Sakr and M.E. El Naggar.1 Geotechnical Research Centre, Faculty of Engineering, The University of Western Ontario, London, ON N6A 5B9, Canada.

M. Nehdi. Department of Civil and Environmental Engineering, The University of Western Ontario, London, ON N6A 5B9, Canada. 1Corresponding author (e-mail: naggar@uwo.ca).

Can. Geotech. J. 41: 313–325 (2004) doi: 10.1139/T03-089 © 2004 NRC Canada

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Research significance

The outcome of this research could be very useful to the construction industry, since implementing fiberglass-reinforced polymer (FRP) composites in deep foundations could result in much longer service life (expected to be three times that of conventional pile materials), which in turn could result in substantial life cycle savings in project costs. FRP piles can also eliminate many durability problems of deep foundations located in harsh environments, such as marine borer attack on wooden piles and corrosion of steel reinforcement in concrete or steel casing piles. Self-consolidating concrete (SCC) can eliminate the lack of cross-sectional integrity and soil encroachments often experienced in deep foundation construction due to the lack of visibility and accessibility.

A new pile driving technique at the pile toe is examined in this study. It can result in substantial improvements in the pile installation process. It allows driving FRP piles in difficult soil conditions such as dense sand. Moreover, when used with conventional steel shell piles, this technique allows reducing the pile wall thickness and improving the driving efficiency. This study also provides designers and contractors with a demonstration case and a database on the feasibility and advantages of FRP–SCC composite piles.

Introduction

Pile foundations are generally used to support structural loads in situations where shallow foundations cannot provide the required bearing capacity or where soil settlement is a major concern. Typical pile materials for deep foundations include steel, concrete, and wood. Wooden piles, precast or reinforced concrete piles, cast-in-place concrete piles, and steel piles have been used in practice for a long time. These pile materials have limited service life, however, and are associated with high maintenance costs when used in harsh environments, for instance because of corrosion degradation and marine borer attack (Lampo et al. 1998). Moreover, lack of cross-sectional integrity through the formation of air pockets and soil encroachments is often experienced in the construction of drilled shafts due to the lack of visibility and accessibility.

A relatively recent trend in deep foundation design is using FRP composite materials in piles because of their light weight, high specific strength, corrosion resistance, chemical and environmental resistance, and low maintenance cost. Composite FRP piling has been used in practice in waterfront barriers, fender piles, and bearing piles for light structures (Iskander and Hassan 1998). Most composite piling products are made of FRPs or high-density polyethylene (HDPE), with fiberglass reinforcement and additives to improve mechanical properties, durability, and ultraviolet light protection. However, FRP composite piles have not yet gained wide acceptance in the deep foundation industry because of the lack of proper design guidelines for predicting their drivability and load-carrying capacity, and the need for construction projects that demonstrate the concept. Implementing the use of FRP materials in piling applications may result in substantial

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benefits to the deep foundation industry, which is the underlying motivation for the research reported herein.

Few studies are available in the literature on FRP piles, probably due to their novelty. Iskander and Hassan (1998) reviewed the available FRP composite products for fender applications. Frost and Han (1999) studied the interface friction between FRP and sand and showed that the interface friction angle between sand and FRP depends on the normal stress level, the shape of sand particles, and the relative roughness of FRP.

Iskander et al. (2001) conducted a parametric study using wave equation analysis on FRP piles. They concluded that the drivability of composite materials depends on the specific weight and elastic modulus of the composite section. Ashford and Jakrapiyanun (2001) compared the drivability of FRP composite piles and steel and concrete piles also using wave equation analysis. They found that FRP piles compared favorably with steel pipes and precast prestressed concrete piles. Mirmiran et al. (2002) conducted a field installation of empty FRP tubes, FRP–concrete piles, and spliced tubes using the conventional top driving technique and concluded that empty FRP piles cannot be driven to great depths and cannot be driven in hard soils such as dense sand. Han and Frost (1999) examined buckling loads of FRP piles during installation and service life. Their study showed that critical buckling loads of FRP piles depend on the shear effect coefficient, the lateral soil resistance, the embedment ratio, the overall boundary conditions, and the critical length. Shear deformation plays an important role in reducing critical loads for FRP piles and steel piles. Critical loads increase with an increase in the shear effect coefficient, lateral soil resistance, and embedment ratio. Han and Frost showed also that the skin friction between the pile and soil plays a very limited role. They concluded that buckling of FRP piles is unlikely to happen except for very long piles or for piles in very soft soils.

Pando et al. (2000) performed a full-scale pile load test at the Route 40 bridge (Nottoway River, Virginia) using FRP tubes filled with concrete and prestressed concrete piles. Their experimental program consisted of pile driving analyses, Statnamic axial compression load tests, and Statnamic lateral load tests on both piles. They found that both pile materials exhibited a similar response with respect to their drivability and axial capacity in compression. The response of FRP composite piles for lateral loads was significantly softer than that of the prestressed concrete piles, however.

Objectives

The primary objectives of this research are to develop an efficient technique for driving FRP piles and thin-walled piles; to experimentally evaluate the performance of FRP–SCC hybrid piles under compressive, uplift, and lateral loading; and to compare the behaviour of FRP–SCC hybrid piles with that of conventional steel piles. The drivability of FRP tubes and the behaviour of FRP–SCC hybrid piles under various loading conditions were simulated in a pressure chamber specially designed to simulate field vertical and radial soil confinement pressures. The performance of FRP pile segments was compared to that of steel pile segments driven using the toe driving technique and also using conventional

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driving at the pile head. The results of this extensive research program are reported in the following sections.

Soil samples for pile installation

Dense sand samples were prepared in a pressure chamber to represent difficult field piling conditions. Fanshawe brick sand was used in the tests and consisted of fine, subround to round, air-dried sand. The grain-size distribution for this sand is shown in Fig. 1. The sand used was classified as poorly graded, with particle sizes in the range of 0.075– 2.00 mm. A standard density test showed that the sand had a maximum unit weight of 17.72 kN/m3 and a minimum unit weight of 14.66 kN/m3. Soil samples for pile installation were prepared using a raining (pluviation) technique, which is known to provide uniform soil properties. The relative density of the soil samples used for this testing program was about 90%, with a standard deviation of 2.5%. Table 1 provides details of the soil properties. The density of the soil samples was monitored using three different techniques to ensure that soil properties do not vary from one pile testing to another: (i) three small soil samples of 106 × 10–6 m3 were taken every 100 mm of soil deposition at three different locations in the pressure chamber to monitor the uniformity of the relative density of the sand in the pressure chamber; (ii) a relatively large sample using a 944 × 10–6 m3 mould was taken at each 100 mm of soil deposition, and at each 200 mm of soil deposition the raining process was stopped and in situ nuclear density measurements were conducted to determine the density and water content of the soil; and (iii) the pressure chamber was weighed both empty and filled with soil to determine the average density of the soil sample. These measurements ensured that a uniform soil sample would be achieved with the desired relative density. Figure 2 shows the density versus depth for the soil samples. It is clear from Fig. 2 that the soil density was reasonably uniform. The disparity between the density measurements obtained from the mould and those obtained from small samples was about 2%, and the disparity between the mould and nuclear measurements was less than 10%. The disparity between unit weight values measured using containers and those from nuclear density measurements can be attributed to decaying radioisotope used in the nuclear density measurements.

Test setup

A large-scale modeling facility was developed at the University of Western Ontario to test large-scale model pile segments of 1.5 m in length at different confining pressures. The key advantage of using large-scale pile modeling over other smaller scale model pile and centrifuge testing is the ability to reasonably simulate vertical and radial effective stress profiles at any depth using a pile with relatively large dimensions. Modeling the pile thickness is of prime importance for FRP composite piles, since these composites are often nonhomogeneous and exhibit anisotropic viscoelastic behaviour (which is pertinent for piles subjected to large lateral loads); a small model pile with a small thickness such as in centrifuge testing may not adequately represent the material nonhomogeneity. Another advantage of using large-scale model piles is the ability to correctly present pile–soil interface characteristics. Moreover, model piles can be easily installed using different

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driving techniques. Lastly, a pressure chamber can be operated and maintained at a fraction of the cost of a centrifuge (Yazdanbod et al. 1984). However, the only limitation for large-scale pile modeling is that it is only applicable to pile segments. Therefore, special caution should be exercised when extrapolating the results of these tests to full-scale piles.

The pile installation setup consists of a pressure chamber, pile driver, toe driving device, model piles, and instrumentation and data acquisition system. Six tests were conducted on FRP tubes, FRP–SCC composites, and steel piles. Figure 3a shows a schematic view of the testing setup, and Fig. 3b shows an oblique view of the pile driving assembly.

Pressure chamber

As shown in Fig. 3a, the soil column is formed in a containment steel cylinder of 1.34 m inside diameter and 1.52 m height. Vertical and radial stresses within the in situ soil at different depths were modeled using vertical and circumferential air bladders with independent control. Therefore, it was possible to establish the ratio of horizontal to vertical effective stresses in the chamber, K0, to be equal to the ambient stress ratio at that depth in situ. A thermoplastic polyfin membrane sleeve was used as a radial pressure bladder. Pressure was applied to the membrane by compressed air through an input port and monitored by a pressure sensor connected to the compressed air supply and two earth pressure cells embedded vertically and horizontally at a depth of 0.6 m to measure the actual radial and vertical pressures in the soil sample. The vertical pressure was applied by a flat gum rubber membrane 4.76 mm in thickness, situated at the top of the chamber. When pressurized, this membrane reacts against a 19 mm steel cover plate. Three ports pass through both the plate and the membrane: a central port with a diameter of 279 mm to accommodate the pile and another two ports with a diameter of 50.8 mm to serve as exit points for an air pressure hose, drainage hoses, and earth pressure cell wires. The rubber sleeve was closed at the bottom by a 50.8 mm thick waffle-type neoprene energy absorber to reduce the energy in waves reflected from the base of the chamber. The rubber tube is impermeable and provides water tightness to the soil column. Field-like conditions were achieved in the pressure chamber by limiting the size of pile specimens to be tested to 168 mm in diameter and installing the pile to a maximum depth of 1.20 m to satisfy a horizontal boundary of eight pile radii, and three vertical radii from the pile tip (Vipulanandan et al. 1989).

Model pile specimens

Two instrumented FRP cylindrical tubes and one steel pile were used in the study. The piles were provided with an interior annular shoe at the pile toe to facilitate their installation using a toe-driving device. The cylindrical steel pile was an open-ended model made of cold-drawn steel tubing 168 mm in diameter with a modulus of elasticity of 2.15 × 105 MPa, a thickness of 6.35 mm, and a total length of 1.524 m. The geometrical properties of the piles are given in Table 2. The FRP composite pile considered in this study was a commercially available FRP pipe fabricated using six layers of glass filament wound at ply angles of 55° and –55°. The glass fiber was

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impregnated with laminating resin at a volumetric ratio of 60/40. The FRP pile shell has a tensile strength of 440 MPa and an elastic modulus of 17 × 103 MPa.

The FRP tubes were filled with a cost-effective, self-consolidating concrete (SCC) specifically developed by Nehdi et al. (2003) for deep foundation applications. The SCC is a material that flows under gravity without the need for mechanical vibration and thus assures the structural integrity of piles and eliminates air pockets and soil encroachments. The SCC mixture proportions (in kg/m3) are as follows: Portland cement ASTM type I (200), ground granulated blast-furnace slag (200), water (150), coarse aggregate with a maximum particle size of 9.5 mm (850), fine aggregate (850), and water-reducing admixture (3.0 L/m3). Fresh concrete properties were as follows: unit weight of 2420 kg/m3 and slump flow of 550 mm. The 7 day and 28 day compressive strength of the SCC placed in the piles was 42 and 58 MPa, respectively.

Pile specimens were instrumented to measure dynamic data during pile driving along with the axial and bending stresses during subsequent static axial and lateral loading tests. A head strain gauge circuit, toe load cell, laser displacement transducer at the pile head, and two piezoelectric accelerometers were used to acquire data during pile driving. Two accelerometers were fixed externally at 200 mm from the pile head along the pile diameter at two opposite directions to measure the vertical acceleration at the pile head.

The instrumentation of the piles for static axial and lateral loading tests consisted of seven levels of strain gauges attached to the external surface of the model piles to provide additional information on the shaft friction along the embedded length of the pile. The strain gauges at each level were wired in two different ways to measure the axial load during compression and uplift loading and the bending stresses during lateral loading tests. The strain gauges were distributed over the length of the piles such that the first bridge was approximately 206 mm from the pile head above the sand surface and the remaining bridges were distributed evenly along the pile length. Protection of the instrumentation was achieved by applying three layers of a coating system, which consisted of a polyurethane gauge top coating to protect the strain gauges from exposure to moisture, a layer of a high shear strength epoxy (M-Bond AE-10), and a fiberglass cloth immersed in a two-part epoxy resin to protect the instrumentation against abrasion due to high stresses developed during driving and subsequent static loading tests. A toe closure device instrumented with a load cell was used to close the pile toe and to measure the pile toe resistance during axial static load tests. Figure 4 shows the details of the toe closure device used for both the steel pile (Fig. 4a) and the FRP composite pile (Fig. 4b).

To facilitate pile installation using toe driving (as discussed later in the paper), a special annular steel shoe was welded (for steel piles) or bonded (for FRP piles) to the pile toe as shown in Fig. 4. The design of the annular steel shoe should be given special care because it is a key element in successful pile installation using toe driving. The exterior diameter of the annular shoe at the pile toe is slightly smaller than the interior diameter of the pile and its height should be where hb is the height of the annular shoe, Fp is the peak impact force, di is the interior pile diameter at the toe, and tb is the bond shear strength.

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The annular shoe was bonded to the toe of the FRP pile using a two-part methacrylate structural adhesive Plexus MA300 (ITW Plexus, Danvers, Mass.). When the fixture time (curing time) of 15 min is achieved, Plexus MA300 develops a bond strength of 21 000 kPa. The shoe length for the current experiment calculated using eq. [1] was 75 mm.

The clearance between the pile and the shoe should not exceed 1 mm, and the surface should be cleaned before adhesive application to achieve high-quality bonding. The application of Plexus adhesive involved three steps: distribute spacers (steel wire) around the shoe with a thickness of 0.8 mm, clean the surface of the shoe and pile using Plexus TC120 primer, and apply the Plexus MA300 to both surfaces and slide the shoe into its location inside the pile toe.

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Pile driving techniques

Driven piles are generally installed by striking a hammer at the pile head. A pile driving system usually consists of a pile, pile cushion, pile cap, hammer cushion (cap block), and hammer (ram). A pile cushion of 51 mm thick plywood was used in this testing program to reduce the peak impact force and the risk of pile head damage.

An innovative technique for driving a tubular pile at its toe using an impact hammer inside the pile, called down-the-hole driving technique, was developed and tested in France (Benamar 2000). The main difference between the conventional driving technique and this technique is that head driving pushes the pile into the soil (compression stress wave), whereas the toe driving technique pulls the pile down (tension stress wave). A pile driving system using the toe driving technique consists of a pile with a special annular shoe at the pile toe, toe driving device, long follower (anvil), hammer cushion, and hammer. The next section describes the principles of the toe driving technique used in this study and outlines its advantages and limitations.

Principle of toe driving

A toe driving system was designed and built at the University of Western Ontario. It involves an independent penetration of the pile toe and the pile itself using a special driving head. The driving head is connected to the hammer through a long steel follower (anvil). Figure 5a shows details of the driving hammer, and Fig. 5b shows an oblique view of the new driving device. The ram of the hammer hits the anvil (1) that transfers the impact shock to a loading plate (3) without any cushion. The loading plate is connected to a conical tip (9) to facilitate pushing the soil away. The tip can be conical with an angle of 60° or adjusted up to 180° (flat bottom). The central part (6) with the conical tip penetrates the soil independently from the pile (7). The pile (7) then follows the central part via a special spring element (4) and a transmitting ring (5). The steel-to-steel contact between the ram and the central part causes a high peak force needed for the penetration in hard soils such as dense sand, boulders, or sandstone layers. The spring transmits the force gradually to the pile toe (7) to pull the pile downwards. The advantage of the spring element is that it generates less noise because there is no steel-to-steel contact between the hammer and the pile wall, and the soil also dampens the noise, another advantage compared to conventional driving at the pile head. A special seal (O-ring) prevents the entrainment of soil into the pile. The new driving technique also enables the use of long hammers inside the pile.

Pile installation using toe driving involves two steps. First, the hammer strikes the pile cushion, the follower transfers the energy to the conical tip of the driving device, and the tip penetrates the soil and compresses the spring element. Second, the transmitting ring is moved downward, a tensile stress wave is generated along the pile shaft, associated with pile elongation and reduction of its cross section, and the pile is pulled downward. The main difference between toe driving and head driving is that in the toe driving the toe resistance is mobilized first and the initial tension wave is strongly attenuated as it travels

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upwards to the pile head. On the other hand, in classical driving, the shaft friction is mobilized first in the pile–soil interface, leading to subsequent disturbance of the surrounding soil. The initial compression wave travels downwards to the pile toe. The soil layer is then compressed, creating a high stress field around the pile and an increase in the shaft resistance.

Many advantages result from using toe driving. First, in the case of FRP composite piles where there is the risk of buckling and the compressive strength of the FRP composite is much lower than its tensile strength, toe driving is considered a very effective option because it utilizes the high tensile strength of the FRP composite during installation. Second, the wall thickness of conventional steel piles can be reduced and the driving efficiency can be considerably improved by using toe driving, since most of the impact load is transferred directly to the soil (Arentsen et al. 1996). Third, open-end piles can be installed using a special recoverable driving head connected to the hammer, therefore accomplishing an economical installation of such tubular piles. Fourth, the pile directional stability during toe driving is improved significantly, since the centre of gravity of the pile and hammer system is much lower. Fifth, piles can be driven through gravel or hard strata such as scattered boulders or very resistant ground, since the energy is concentrated at the pile toe. The toe driving is particularly suitable for urban areas because less noise is generated. Lastly, using this technique leads to a significant reduction in driving time and a higher load bearing capacity (Benamar 1999). The limitation to the new technique is that the driving efficiency is less obvious under easy driving conditions.

Test procedure

Six pile driving tests were conducted, two using driving at the pile head and four using toe driving. Several criteria were considered to develop a representative modeling of pile driving in a laboratory setup, including (i) ambient effective stresses in the 1.34 m diameter pressure chamber with vertical pressures of 60 and 120 kPa and corresponding radial pressures of 30 and 60 kPa were maintained during the driving process and subsequent axial load testing to simulate the vertical and horizontal stresses in normally consolidated dry sands (K0 = 0.5) at depths of 4.0 and 8.0 m (24 and 48 pile diameters); (ii) the pile penetration per blow was maintained in the normal range achieved during in situ pile driving (2.5– 5.0 mm/blow); and (iii) the surface texture of the model piles was similar to that of typical full-scale steel pipe and FRP composite piles.

Each of the six tests involved the following steps: (i) pile driving, (ii) uplift tests, (iii) pouring concrete followed by 7 days of curing, (iv) axial compression–uplift–compression test, and (v) lateral loading test. Five tests were conducted using vertical and radial pressure combinations of 60 and 30 kPa, and the sixth test was conducted using vertical and radial pressure combinations of 120 and 60 kPa. These different vertical and radial pressures were applied to the soil sample in the pressure chamber to simulate a 1.2 m long pile segment at 4.0 and 8.0 m depths in normally consolidated sand (K0 = 0.5) assuming average unit weight of the soil (. =16 kPa). After ensuring that stresses within the soil reached equilibrium, the pile was driven usinga1kN weight falling from a height

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of 1.2 m at a rate of 6 blows/min until it reached a final penetration depth of 1.2 m. Immediately after driving, two uplift tests were conducted on the piles up to a displacement of 0.15 pile diameters at vertical and radial pressure combinations of 60 and 30 kPa and 120 and 60 kPa, respectively. The pile was then redriven to its initial position, a toe closure mechanism was lowered into the pile toe and placed in its location, and the FRP pile was then filled with SCC. A curing time of 7 days was allowed. A compressive axial loading test to a vertical displacement of 0.20 pile diameters was then conducted followed by an uplift loading and a compressive test. Lastly, a lateral loading test was applied to the pile until a final displacement of 0.10 pile diameters. Results and discussion

Pile installation

As described earlier, FRP piles were driven empty into the soil using toe driving and then filled with concrete and tested. Another driving process using head driving was conducted. In the toe driving installation, two driving energies of 0.666 and 1.200 kJ using a single-acting hammer with a falling height of 1.2 m and hammer weights of 0.55 and 1.00 kN were used. Figure 6a shows a comparison of driving records for the two different driving energies in terms of toe penetration (z) normalized by pile diameter (d). It can be observed that for the same penetration, the 0.55 kN hammer required on average twice the number of blows of that required using the 1 kN hammer. Therefore, it was decided to use the 1 kN hammer in the rest of the tests to minimize the time required for driving the piles and to achieve a penetration rate of about 2.5 mm/blow.

Figure 6b shows a comparison of the average penetration resistance measured for two different confining pressures. Vertical and radial pressure combinations of 120 and 60 kPa and 60 and 30 kPa were used, which correspond to soil penetration depths of 8.0 and 4.0 m, respectively, in normally consolidated sand (K0 = 0.50). On average, at the greater soil depth, 75% more hammer blows were required to achieve the same pile penetration depth. Figure 6b also shows that the blow count was much higher at the lower four pile diameters as a result of sand densification around the driven pile. It can also be noted from Fig. 6 that the blow count was almost constant at the last four pile diameters, as indicated by the constant slope of the cumulative hammer blows versus depth ratio. This shows the weak influence of shaft friction on the penetration at greater depths. This is because the toe driving mechanism transmits most of the hammer blow energy to the soil at the pile toe directly, creating a cavity below the pile toe, and only a fraction of the blow energy is transferred to the pile shaft, pulling it down into the created cavity. The tensile force acting on the pile during impact elongates the pile and reduces its cross-sectional area. The reduction in the pile cross section decreases the lateral soil pressure at the pile–soil interface during driving and therefore reduces the shaft resistance slightly.

A steel pile was installed using toe driving and then closed at the pile toe. Another installation of the steel pile using head driving was conducted. Figure 7 shows a comparison between FRP and steel piles installed using toe and head driving techniques.

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In general, the FRP pile installed using toe driving required about 43% less energy than the FRP–SCC pile installed using conventional head driving. The much easier driving of the empty FRP pile using toe driving, despite its low impedance, can be explained by the role of the driving mechanism, which applies the impact directly to the soil and increases the overall impedance of the pile-driving system. Comparing the steel and FRP pile installation using toe driving showed that the steel pile required about 7% less energy than the FRP pile.

It can be also noted that for the steel pile, the driving record was about the same regardless of the installation method. This is likely a result of the high impedance of the steel pile. It is worth mentioning that no signs of damage were observed during driving empty FRP tubes, even though six driving processes were conducted on the same pile. FRP–SCC hybrid piles were also installed using head driving without showing any visible cracks or damage (see Fig. 8).

Stress wave propagation in piles

Typical time–force signals at the pile head and near the pile toe recorded during toe and head driving for piles installed at vertical and radial pressures of 60 and 30 kPa are shown in Figs. 9 and 10 for steel and FRP piles, respectively. Figures 9 and 10 show that in the toe driving installation the initial stress wave at the pile toe is tensile and propagates upwards from the pile toe to the pile head. The tensile wave is then reflected at the pile head as a compressive wave. After the first up and down cycle of travel of the stress wave in the pile, its amplitude is substantially reduced by shaft and toe interaction with the surrounding soil. The initial tensile wave first interacts with the soil near the pile toe. The reflected waves from the pile toe are compressive stresses and superpose with the initial tensile wave induced by the ram stroke. The up and down travel time of the wave is measured for FRP empty pipe, steel, and FRP–SCC composite piles and agrees with the theoretical value of 2L/C, where L is the pile length and C is the wave velocity (3260 m/s for FRP empty pipe, 3730 m/s for FRP–SCC composite pile, and 5200 m/s for steel pile).

In head driving (Figs. 9, 10), the initial wave at the pile head is compressive and propagates to the pile toe. The extent of the stress wave reflection at a return time of 2L/C is characterized by a tensile dip in the force near the pile head. Figures 9 and 10 also show that there is a distinct difference between the force amplitude of piles installed using head driving and that of piles installed using toe driving, since in the latter most of the impact load is transferred directly to the soil and part of the impact is transferred to the pile shaft to pull the pile downwards. For example, the peak force measured for steel piles driven using head driving was 290 kN (compression) compared with 116 kN (tension) for piles installed using toe driving, and similarly the peak forces for FRP composite piles installed using head and toe driving were 282 kN (compression) and 125 kN (tension),

Fig. 9. Typical stress wave data for steel cylindrical pile: (a)at Fig. 10. Typical stress wave data for FRP cylindrical pile: (a)at pile head; (b) near pile toe. pile head; (b) near pile toe. respectively. The difference in force amplitudes (which is about 60%) is

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transferred to the soil directly at the pile toe in toe driving, and only 40% of the force is transferred to the pile material.

Figure 11 schematically shows the main difference between wave propagation in the pile during toe driving and that during head driving. In conventional driving at the pile head, the shaft friction is mobilized first and the interaction between the pile–soil system leads to subsequent disturbance of the surrounding soil. When the initial compressive wave travels downwards to the pile toe (Fig. 11a), the soil layer is compressed, creating a high-stress field around the pile and an increase in the shaft resistance is observed. In toe driving, the toe resistance is mobilized first and the initial tensile wave is strongly attenuated as it travels upwards to the pile head (Fig. 11b). The energy transfer from the pile to the soil is made through the tensile wave transmitted and reflected at the pile toe.

Static resistance of piles

Pile axial compression tests Pile specimens were tested in compression at two different vertical and radial confinement pressure combinations of 60 and 30 kPa and 120 and 60 kPa, respectively. The compressive bearing capacity of piles, Qc, was taken as the load required for displacing the pile head by 10% of the pile diameter (De Nicola and Randolph 1999). Figure 12 shows the load–displacement curves for steel and FRP piles installed using toe and head driving. Figure 12 shows that for all compression tests there is a clear tendency for the bearing capacity to continue to increase with increasing pile dis-placement up to a displacement ratio of about 0.2. Piles installed using toe driving showed a similar response at low displacement levels and a stiffer response at high displacement levels (>7% of pile diameter) compared with piles driven at the head. For example, an FRP–SCC pile installed using toe driving (Fig. 12a) and tested at vertical and radial confinement pressure combinations of 120 and 60 kPa and 60 and 30 kPa showed an increase in the load bearing resistance of up to 18% and 10% compared with an identical pile installed using head driving. A steel pile installed using toe driving (Fig. 12b) showed a similar response to that of an identical pile installed using head driving at low displacement levels and a slightly stiffer response at higher displacement levels. This can be explained by the desirable effect of toe driving of densifying the soil around the pile toe, therefore increasing the pile resistance. Comparing Fig. 12a (FRP–SCC pile) and Fig. 12b (steel pile) reveals that both piles had a similar response in compression.

Table 3 shows the axial capacity of the various piles tested. The FRP–SCC piles showed resistance comparable to that of the steel piles. The end bearing resistance of piles driven using different techniques ranged from 48% to 53% of the axial capacity. The measured values of bearing capacity were generally about 30% lower than values reported in the literature for piles driven in dense sand (Sakr et al. 2004). This could be a result of using a neoprene energy absorber with high compressibility at the bottom of the pressure chamber.

Pile uplift tests

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For each pile, an uplift loading test was conducted immediately after driving at a vertical and radial confining pressure combination of 60 and 30 kPa. The pile was then redriven to its initial position and a second uplift loading test was conducted at a vertical and radial confinement pressure combination of 120 and 60 kPa. The results of the uplift loading tests for FRP–SCC hybrid piles, FRP piles, and steel pipes under various confinement pressures are plotted in Fig. 13 in terms of the load applied at the pile head versus the displacement ratio, which is defined as the ratio of the pile head upward movement to the average pile diameter. Figure 13 shows that the FRP empty tube installed using toe driving had a response similar to that of the FRP–SCC hybrid pile, regardless of the installation method. The peak tensile loads were achieved within an approximate displacement of 10% of the pile diameter. FRP–SCC and FRP empty piles tested at low confining pressures (60 and 30 kPa) showed similar initial stiffnesses at the early stages of loading (up to 1.5% of pile diameter), whereas at high pressures the hybrid FRP–SCC piles showed a stiffer response, especially at higher displacement ratios. The uplift capacity, Qup, is defined as the uplift load that corresponds to a pile head displacement of 5% of the pile diameter (Kulhawy and Hirany 1989). Table 4 shows the uplift capacity, Qup,of the FRP and steel cylindrical piles installed using toe and head driving and tested at different testing confinement pressures. Table 4 shows that the FRP piles yielded an uplift capacity similar to that of the steel piles, regardless of the installation method. The uplift capacity ratio, KV, was used for the comparison of uplift capacity due to different installation methods. The uplift capacity ratio is defined as the ratio of the uplift capacity of the pile installed using toe driving at a certain confining pressure to the uplift capacity of the same pile installed using head driving at the same pressure. The KV values ranged from 0.9 to 1.0, meaning that toe driving had a minor effect on the pile performance during subsequent static load testing.

Pile lateral load tests The lateral load tests of piles were carried out at vertical and radial confining pressures of 60 and 30 kPa, respectively. Figure 14 shows the load–displacement curves for cylindrical steel and FRP–SCC piles driven using toe and head driving and then laterally loaded up to a displacement of 0.1 pile diameters at the pile head. Figure 14 shows that the response of all piles tested was approximately linear, regardless of the installation method. The response of piles installed using toe driving was very similar to that of piles installed using head driving. As expected, the response of steel piles was stiffer than that of the FRP–SCC piles as a result of the higher flexural stiffness of steel piles compared with FRP–SCC piles. A similar observation regarding lateral resistance of FRP composite piles (low flexural stiffness) compared with that of prestressed concrete piles (high flexural stiffness) was reported by Pando et al. (2000). Their field tests on piles subjected to Statnamic lateral loads showed that prestressed concrete piles had higher lateral resistance than FRP composite piles. The ultimate lateral load is determined based on the lateral load applied at the pile head at a lateral pile head displacement of 6.25 mm (Prakash and Sharma 1990). Table 5 shows the ultimate lateral load for the tested piles. Table 5 shows that the ultimate lateral load of FRP–SCC and steel piles installed using toe driving was comparable to that of piles installed using head driving. For example, the ultimate lateral load of the FRP–SCC pile installed using toe driving was 8.35 kN and that for the FRP–SCC pile installed using head driving was 8.25 kN. The ultimate lateral

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loads of the steel pile installed using toe and head driving processes were similar and about twice that of FRP–SCC piles due to the superior flexural rigidity of steel piles (see Table 5).

Guidelines for toe driving and pile design

The current study presents a novel toe driving technique suitable for the installation of piles in hard soil conditions and demonstrates the effectiveness of this new technology. FRP composite shells were installed using the new driving technique, utilizing 60% of the driving energy required for the conventional head driving in dense sand. Based on the observations made during this laboratory testing program, the following aspects should be taken into consideration for the application of the new technology for pile installation:

(1) A preliminary drivability study is required to choose a suitable hammer and to predict the driving energy and maximum forces developed during the impact. Different toe driving components including the anvil, loading plate, central part, and steel base should be designed to safely withstand and transfer the peak forces without yield. For long prototype piles, buckling should be considered in the anvil design. A factor of safety of 2 was considered in the design of different components of the device used in this study. Due to the repetitive nature of impact loads, however, fatigue should be considered by applying a higher factor of safety in the design of the prototype toe driving device.

(2) The spring elements and transmitting ring are responsible for pulling the pile shaft down, and the pile shaft resistance governs their design. These components are thus designed using the dynamic shaft resistance (from the drivability study) and applying a factor of safety of 3. The factored shaft resistance is used to determine the stiffness of the spring element, the strength of the transmitting ring, and the contact area between the transmitting ring and pile toe.

(3) To facilitate pile installation using the toe driving technique, an annular steel shoe should be welded (for steel piles) or bonded (for FRP composite piles) to the pile toe. The height of the steel shoe should be designed so that the bond strength of the connection exceeds the factored dynamic shaft resistance. The thickness of the annular shoe ring is also designed to withstand the factored dynamic shaft resistance.

(4) Considering only the dynamic shaft resistance, which could be considerably less than the total resistance, the thickness of a steel pipe pile installed using toe driving can be optimized. For example, the current tests showed that the peak driving force during toe driving was about 40% of that force in the case of head driving. This means that a reduction of about 60% of the pipe thickness could be achieved when using toe driving. Moreover, toe driving provides a viable option for driving empty FRP tubes, which could not be installed using conventional head driving, as observed by Mirmiran et al. (2002).

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(5) The effectiveness of the toe driving device is more pronounced in soils where toe resistance represents a high proportion of the total capacity, such as in the case of piles installed in sand. For soil conditions where the toe resistance is 20% or less of the total capacity, e.g., piles installed in clay, the new device may not be provide additional efficiency over conventional driving.

Conclusions

An experimental investigation of the drivability of FRP– SCC composite piles using a new toe driving technique and conventional head driving in dense sand was presented in this paper. Simulating prototype piles driven to different penetration depths ranging from 4.0 to 8.0 m (24d to 48d) was achieved using a specially designed pressure chamber. An innovative technique for FRP cylindrical pile installation was designed and tested in this study. FRP empty piles and steel piles were installed using this new technique to demonstrate its usefulness compared to conventional head driving.

The new toe driving method offers a viable driving option for the installation of empty FRP and thin-walled steel piles. Moreover, toe driving proved to be efficient in hard soils such as dense sand.

The final penetration depth was achieved without any damage or overstress in the pile because of the strong retrievable pile point.

The FRP pipe is filled with SCC after driving to increase its structural resistance to axial and lateral loads. The durability and corrosion resistance of FRP–SCC composite piles can substantially increase the lifetime of deep foundations and therefore reduce life-cycle costs.

Static load testing of steel and FRP–SCC piles in compression showed that the composite piles yielded a response similar to that of steel piles. Piles installed using toe driving showed a stiffer response than that of piles installed using head driving as a result of densification of the sand around the conical tip. The axial uplift capacities of empty FRP, FRP–SCC, and steel piles driven at pressures of 60 and 30 kPa and tested at different confinement pressures were comparable. Pile installation using toe driving showed a minor effect on the uplift capacity of steel and FRP composite piles. Lateral load tests on piles showed that the ultimate lateral loads of FRP–SCC and steel piles were independent of the driving method. Moreover, the ultimate lateral load of steel piles was higher than that of FRP–SCC piles. This is not normally a major concern, however, since the lateral capacity of piles is usually one order of magnitude lower than their axial capacity. Also, FRP piles can be designed to sustain higher lateral loads by either using a batter or increasing the flexural rigidity of the FRP by using stiffer aramid or carbon fibers, but the cost effectiveness should be evaluated.

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It is anticipated, however, that the current trend in using FRP composite for many civil engineering applications including deep foundations will result in reducing their current high costs.

The findings of this research are based on an experimental investigation of pile segments in a large-scale laboratory facility. Full-scale testing using toe driving in various subsurface conditions is required to confirm the findings, however. In addition, a numerical model using wave equation analysis for FRP composite piles is currently being developed to optimize different parameters and evaluate drivability.

Acknowledgements

The authors would like to thank Mr. Carl Ealy, head of Deep Foundation Research at the U.S. Federal Highway Administration (FHWA), for providing the pile driving hammer and Mr. Helge Wittholz of Polymarin-Bolwll Composites Inc., Huron Park, Ontario, for providing the FRP 0° pipes examined in this study. The authors would also like to thank the reviewers for their valuable comments.

References

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Iskander, M.G., and Hassan, M. 1998. State of the practice review in FRP composite piling. Journal of Composites for Construction, ASCE, 2(3): 116–120.

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List of symbols

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C wave velocityCc curvature coefficientCu uniformity coefficientd average diameter of pile shaftdi interior pile diameter at toeemax, emin maximum and minimum void ratio, respectivelyE elastic modulus hb height of annular shoe K0 coefficient of lateral earth pressure at rest KV uplift capacity ratio L pile length Qc ultimate pile capacity in compression QL lateral capacity Qup ultimate uplift load of cylindrical pile t thickness of pile walltb annular pile shoe thicknessts thickness of the pile wallz embedded pile depth at a certain location dF interface friction angle (sand/FRP) ds residual interface friction angle (sand/steel) f friction angle . unit weight of soil tb bond shear strength

end.

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