Raymond Castelli and Mr. John Wisniewski also of PB Americas, Inc. for guidance during the field program. Opinions expressed in this paper are solely of the ...
Installation Monitoring of Open End Pipe Piles Malhotra, S. PB Americas, Inc., One Penn Plaza, New York, NY-10119, United States of America
Keywords: pipe piles, pile driving, plugging ABSTRACT: This paper presents the results of a pile installation monitoring and performance assessment program for pile foundations supporting 14 piers and 2 abutments of the 1700-ft long bridge crossing over a river in Baltimore, Maryland. Dynamic measurements were conducted during installation and restrike of 16 test piles, one at each support location. The foundation piles were 30-inch diameter, 0.75 inch thick pipe piles at the piers and 24-inch diameter, 0.5 inch thick at the abutments. The piles at the piers were approximately 100 ft long and were designed to carry 260 kips service load with a factor of safety of 2.25. The piles were driven open ended into a variable site stratigraphy containing recent deposits of organic clayey silts near the mud line followed by loose to medium dense sand, with some silt, and then by soil deposits from the Cretaceous period consisting primarily of dense to very dense, poorly- to well graded sands and gravels and cobbles, with occasional seams of very hard to very stiff, low to moderate plasticity silts and clays. This paper presents the approach taken to evaluate the driven index piles and to develop criteria for driving production piles, and provides the Engineer with some data on pile driveability in Cretaceous soils in the Baltimore area, soil set-up effect, and an estimate of pile capacity.
1 INTRODUCTION The typical approach to pile foundation design and construction should include an indicator pile program with dynamic measurements and a static load test program. However, very often paucity of funds and schedule constraints cause an owner to not perform conventional static load testing which can take $25,000 to $100,000 and several weeks per test. This paper examines an alternate (though not equivalent) truncated approach to foundation assessment and presents a case history where this approach was implemented. Dynamic measurements were conducted during pile installation for the 14 piers and two abutments of the new 1700-ft long bridge over a river in Baltimore, Maryland. Strain transducers and accelerometers were attached to the pile top during initial driving and restrike of 16 pipe piles. The foundation piles consisted of 30-inch outside diameter, 0.75-inch thick pipe piles at the piers and 24-inch outside diameter, 0.5 inch thick at the abutments. A total of 230 piles were installed.
The piles were driven open ended at a site with variable subsurface conditions. A site specific subsurface investigation including a soil boring program and a soil testing program in the laboratory was implemented during the preliminary design of the bridge foundations. The soil boring program consisted of a total 23 hollow-stem auger borings designated as B-1 through B-23, ranging in depth from 63 ft to 130 ft, and spaced approximately 70 to 120 ft. The locations of these borings are shown in Figure 1. The laboratory testing program included soil classification and index property tests, gradation analyses, triaxial strength tests and compressibility tests. The approach taken to assess the foundation performance included 1) a pre-field wave equation driveability analyses, 2) field pile driving measurements, 3) post-field Case Pile Wave Equation Analyses (CAPWAP), and, 4) revised wave equation analyses based on field and CAPWAP data. The following sections describe in greater detail each of the steps taken to assess foundation performance.
Figure 1: Bridge Layout and Locations of Borings
Figure 2: Generalized Subsurface Profile 2 SUBSURFACE CONDITIONS
to 66 ft between Pier Nos. 2 and 7.
The bridge site lies within the Coastal Plain Physiographic Province. The Coastal Plain Province is characterized by a gently rolling topography with shallow, open valleys of gravel, sand, and clay deposits of the Pleistocene terrace plains, over older Cretaceous deposits. Cretaceous deposits of coastal eastern North America underlie broad areas of the continental margin as eastward tilted layers of gravel, sand, silt, and clay, which overlap rocks of the eastern Piedmont. Figure 2 presents a generalized subsurface profile for the bridge alignment based on the data obtained from the borings drilled for this investigation. The soil borings indicate that the project site is underlain by four principal soil strata. Stratum 1 consists of fill material composed primarily of brown to gray, loose to medium dense, coarse to fine sand, with some silt and clay, and trace of gravel. Stratum 2, which covers the entire project area, consists of recent deposits of low to high plasticity, organic clayey silts. This stratum is highly compressible. The thickness of this stratum varies along the bridge alignment from approximately 18 ft
Stratum 3 consists of Pleistocene era deposits of medium dense, poorly to well graded sand, with trace to some silt, and occasional beddings of poorly- to well-graded gravels and low plasticity silts. The thickness of this stratum varies along the bridge alignment as well as laterally and ranges from 26 to 43 ft. Stratum 4 represents soil deposits from the Cretaceous period and consists primarily of dense to very dense, poorly to well graded sands and gravels, with trace to some silt and clayey silt. It also contains zones of very stiff, low to moderate plasticity silts and clays. Fairly thick zones (>25 ft) of very stiff clay and silts were encountered between Piers 9 and 12. The N-values for this stratum are generally greater than 50 blows per foot (bpf). The elevation of the top of this stratum varies both longitudinally and laterally, and is deepest at the location of the existing river channel. Two groundwater observation wells installed near the abutments indicated groundwater levels between El. +3.35 and +3.75 with reference to the City datum. The mean high water level in the river is El. +1.4 ft.
Table 1: Summary of Design Soil Parameters Range of Total Unit Angle of Undrained SPT NWeight Internal Shear Values Friction Strength *Su/Po’ Stratum No. (blows/ft) (pcf) (deg.) (ksf) 1 6-10 110 30 2 0-2 97 0.26 3 0-50 34 4 >50 38 3-4 *ratio of undrained shear strength to the effective overburden stress
3 DESIGN SOIL PARAMETERS A summary of the design parameters recommended for the four soil strata described above are presented in Table 1. 4 RIVERBED SCOUR A scour study for the river at the location of the bridge indicated scour depths below existing mudline in the main channel (Piers 3 through 7) as deep as 32 ft and 37 ft for the 100-year and 500-year floods, respectively. Scour depths of about 14 ft and 18.5 ft are estimated for the left overbank (Piers 1 and 2) and, 11 ft and 15 ft are estimated for the right overbank (Piers 9 through 14) for the 100-year and 500-year floods, respectively. Scour is not expected to occur at the abutments. The 500 year scour level is shown on Figure 2. The significance of the scour study is that the lateral and axial pile design would have to account for loss of soil support from scour. However, given the low shear strength of Stratum 1 soils, the loss of soil support has relatively minor impact on foundation design.
2000). Since there was uncertainty with regards to the occurrence of plugging, for purposes of comparison, axial capacity of a fully plugged pile was also computed. The tip capacity for plugged open-ended pipe piles is equal to that of a closedended pile of an equal diameter. Graphical plots of predicted pile capacity versus pile penetration for three locations are presented in Figures 3 a, b and c. It is evident from Figures 3a, b and c that the piles at Pier 2 are unlikely to plug, the piles at Pier 7 might form a temporary plug near El. -87 ft, and the piles at Pier 11 are quite likely to plug below El. -52 ft.
5 AXIAL PILE CAPACITY ANALYSES For the piers, the pile lengths were estimated for a service level load of 260 kips with a factor of safety of 2.25 based on a post-scour soil profile. The static pile capacity calculations were performed using the conventional procedures outlined in NAVFAC DM 7.02 and API-RP2A guidelines which recommend limiting values of skin friction and end resistance. Since the piles were expected to be driven openended, the tip capacity was simply taken as the capacity in end bearing of the annular rim of the open-ended pile. However, the total capacity is increased by the lesser of the total internal shaft friction and the end bearing of the plug (Malhotra,
Figure 3a: Measured and Predicted Axial Pile Capacity for Test Pile at Pier 2.
6
PILE INSTALLATION MONITORING APPROACH
The approach taken to assess the foundation performance is shown schematically in a Process Flow Diagram, Figure 4. It includes: 1) a preliminary wave equation driveability analyses, 2) field pile driving measurements, 3) post-field CAPWAP analyses, and, 4) revised wave equation analyses based on field and CAPWAP data.
Figure 4: Approach for Performance Assessment of Piles Figure 3 b: Measured and Predicted Axial Pile Capacity for Test Pile at Pier 7.
Preliminary Wave Equation Analyses: Onedimensional wave equation analyses have been widely used to evaluate pile driveability and hammer performance. The purpose of the wave equation analyses is to predict the behavior of the pile during installation for specific site conditions and driving equipment. The study estimates driving resistances at various depths during driving (End of Initial Drive, EOID). Estimates of driving resistances for various hammer efficiencies are also obtained. The results are used to evaluate the suitability of the selected hammer in driving the selected pile to the design depth and capacity. Field Dynamic Measurements: Dynamic measurements provide data for assessing the hammer and cushion system, measure the energy transferred to the pile top, and can identify localized stresses that might occur in the pile during driving. Moreover, estimates of pile capacity are obtained from these measurements. Information on set-up is also obtained by comparing dynamic measurements from EOID with those obtained at the beginning of restrike tests.
Figure 3 c: Measured and Predicted Axial Pile Capacity for Test Pile at Pier 11.
Post-field Assessment: These analyses use the actual measured transferred energy and stress wave in the pile. First the CAPWAP program integrates the measured acceleration, with time to obtain velocity. It then applies the velocity time history to a lumped mass model of the pile and computes a
predicted force-time history from the change in momentum of each lumped mass. The CAPWAP user compares the measured and predicted force time histories. The match of the computed and measured force time histories is obtained by trial and error by changing the soil resistance and associated parameters. The soil model parameters that result in a reasonable match between the computed and predicted force time histories are considered to be the most appropriate values. The results of the CAPWAP analyses include the magnitude and distribution along the pile of dynamic resistance forces and a best estimate of the pile capacity. Revised Wave Equation Analyses: Wave Equation analyses calibrated to the actual performance of the test piles are performed to develop pile driving criteria for production piles. 7 PRELIMINARY DRIVEABILITY STUDY The computer program GRLWEAP developed by Goble, Rausche and Likens (GRL) (1997) was selected for the pre-field wave equation analyses performed to 1) estimate the feasibility of driving the piles to design depth with the selected hammer, 2) evaluate pile tensile and compressive stresses and pile driving resistances for the selected range in hammer energies, and, 3) develop preliminary driving criteria. Since the piles were to be driven in the pre-scour soil profile, the design called for a minimum penetration and a required ultimate capacity (Ru). For the piers Ru, taken as the sum of the service load times 2.25 and the shaft resistance anticipated within the upper organic clays and silts, was about 630 kips. Extensive parametric studies have been performed by others (Dover et al, 1982) to investigate the sensitivity of predicted capacity to variation in hammer efficiency, cushion stiffness, damping, quakes, and skin friction resistances. It has been found that side damping is a particularly important parameter. For this study, damping and quakes recommended by GRLWEAP were used. In clays the dynamic soil resistances encountered during driving differ from static soil resistances for several reasons. These include phenomenon such as sensitivity of cohesive soils, cyclic mobility, and Poisson’s effect. For those piles to be driven primarily in sandy soil profiles, the dynamic soil resistance was taken as the same skin friction and end bearing as those computed for static pile capacity using the method recommended by API
(1993). For the piles to be driven primarily in clay soil profiles the shaft resistance between the pile and the clayey silts and clays of Stratum 2 and 3 was assumed to be 50 percent of the static shaft resistance. A Delmag hammer Model D30-32 with a rated energy of 74 kip-ft was selected by the Contractor for the preliminary pile driveability study. The hammer manufacturer recommends a 4-inch thick Micarta/Aluminum hammer cushion and reports an efficiency of 0.8 for the hammer. These values were used in the analyses. The driveability study predicted the stresses and driving resistances that will be encountered during pile driving at Piers 2, 7 and 11 (Figures 5a, 6a and 7a). The results of the wave equation analyses indicated that with the Delmag D30-32 hammer, the piles could be driven to the required design depth (80 to 90 ft below mudline) with a reasonable range of driving resistance (60 to 100 bpf). Moreover, the maximum predicted driving stresses, both compressive and tensile, were below the allowable limit of (0.9 fy) = 31.5 ksi (AASHTO, 1992). For the various piers, the driving resistance required to achieve the required ultimate capacity (Ru) for a range of strokes was developed from the preliminary wave equation analyses for the Delmag D 30-32 hammer. 8
DYNAMIC MEASUREMENTS
Dynamic pile testing using the Pile Driving Analyzer (PDA) was performed on the 16 index piles. Instrumentation for dynamic measurement of the pile included two strain gages and two accelerometers mounted near the top of the pile and connected through a coaxial cable to the PDA equipment which includes a data recorder. Initial Drive Monitoring: One pile at each of the 16 foundation locations was instrumented and monitored during initial driving. These piles have a final penetration of about 80 to 90 ft below the mudline. The measurements for initial driving were conducted immediately after the piles penetrated under their own weight by about 40 ft below the mudline and continually thereafter until final penetration. Measured strains and accelerations were recorded on the PDA. Pile capacity obtained from dynamic measurements during initial driving are also presented in Figures 5b through 7b.
(a)
(b)
Figure 5: Measured and Predicted Soil Resistances at Pier 2.
(a)
(b)
Figure 6: Measured and Predicted Soil Resistances at Pier 7.
(a)
(b)
Figure 7: Measured and Predicted Soil Resistances at Pier 11. Restrike Monitoring: Restrike on each pile was performed using the same driving equipment, cushioning material and thickness, stroke and fuel setting as used at the end of initial driving. The hammer was warmed up before the restrike test. The warm-up procedure consisted of a minimum of 20 blows of the hammer at full stroke on a pile that was at least 25 ft from the pile to be restruck. The restrike test consisted of driving the pile a distance of 6 inches, or to a total of 50 blows, whichever occurred first. The number of hammer blows was recorded for each 1 inch or less of pile penetration. With a few exceptions the restrike test was conducted 5 days after initial driving. The ratio of PDA resistances from the beginning of restrike (BOR) to the end-of-initial driving (EOID) for each test pile is presented in Figure 8. Soil setup is an increase in pile capacity resulting from increase in soil shear strength due to dissipation of excess pore pressure generated during pile driving. Increases in driving resistance were observed during restrike indicating that soil setup on the order of 5 percent to as much as 40 percent occurred at the site. The higher values of setup occurred at Piers 9, 11, 12 and the South abutment. Borings at these locations indicated that the bearing stratum, Stratum 4, consisted of stiff to hard silty clay. Increases of 40 percent in capacity
within a short period of 5 days might not be entirely explained by dissipation of pore pressure in this clayey material. It is conceivable that much of the increase in capacity could have occurred because of plug formation. Any setup that occurred in the upper cohesive soils could not be counted upon since these soils were above the design scour level and would be eroded in the future.
Figure 8: Estimates of Set-up (after 5-days)
9 FORMATION OF SOIL PLUG
10 HAMMER PERFORMANCE
Sometimes, open-ended pipe piles get plugged during driving. Plugging occurs mainly due to the mobilization of friction between the soil core and the internal walls of the pile. By definition, full plugging is said to have occurred when there is no relative movement between the soil plug inside the pile and the pile itself. Fully plugged piles behave as closed-ended, full-displacement piles, which are known to have a higher unit skin friction. Thus an open-ended pile that gets plugged during the course of driving would have a unit skin friction between that of a fully unplugged pile and a closed-ended full-displacement pile. Plugging also has an effect on the ultimate bearing capacity of the piles. It also has an effect on the timedependent capacity of piles and more so in pile groups, where the dissipation of excess pore pressures takes longer. The excess pore pressure inside the pile may take much longer to dissipate, compared to pore pressure outside, thus making matters more complex. The occurrence of plugging also changes the dynamic characteristics of piles, leading to situations where the selected hammer may not be suitable to drive a pile to the design depth. There can be wasteful expenditures involved in obtaining a heavier hammer or jetting the pile to the design depth, leading to uncertainty in the actual capacity. To add to the uncertainty, a previously plugged pile may become unplugged during further driving (Randolph, 1990). This slip may occur due to the development of high pore pressures in the soil inside the pile or from “Poisson’s Effect.” Poisson’s effect can occur, when a stress wave from hammer impact travels down the pile causing the pile to expand and subsequently slip relative to the inner soil. At present the prediction of plugging has not been established. Currently available plug prediction methods are mostly based on laboratory model tests (Randolph, 1985, Paikowsky, 1990). However, comparison of total internal shaft friction with plug end bearing can be used as an indicator of whether plugging will occur. To ascertain whether a pile is plugged or unplugged during driving, soil plug measurements should be taken just before and after restriking the pile. However, often field constraints such as lack of accessibility or schedule driven constraints, prevent measuring soil plug heights, which was the case at the subject site.
Hammer energy was monitored for various test pile installations. Energy transmitted from the fully warmed up hammer towards the end of driving the test pile at Pier 2 is shown in Figure 9a. Also shown is the energy transmitted from the same hammer while redriving the same pile the following day. The energy transmitted shown is from the hammer blows within one foot of pile penetration. Moreover, the measured soil resistances did not vary much, thus allowing us to make this comparison. It was found that hammer energy output slowly becomes uniform as the hammer progressively warms up during the driving operation. The warmed up hammer indicates a more consistent energy output. It can be shown that erratic hammer energy leads to an unreliable estimate of pile capacity if one bases the prediction on blow counts alone.
(a) Transmitted Energy (kip-ft) 60 50 40 30 20
Post Maintenance
10
Pre-Maintenance
0 5
6
7
8
9
10
11
Stroke (ft)
(b)
Figure 9: Hammer Performance Before and After a) Warmup, and b) Maintenance Efficiency of a diesel hammer can vary considerably (33% to 80%) because it depends not only on the make and model of the hammer but also on the level of maintenance. Therefore, an accurate estimate of energy transmitted into the pile is important. Moreover, during production driving, sometimes the characteristics of the pile driving system may change, consequently
invalidating the driving criteria. Sometimes, a simple maintenance of the hammer can change the driving energy substantially. For example, after driving the first test pile at Pier 2, the contractor dismantled the open ended diesel hammer, inspected the compression rings and replaced the fuel pipe which increased the energy from the hammer by as much as 50 percent (Figure 9b). In view of the above reasons, to ascertain hammer performance during production pile driving and to validate the driving criteria, PDA testing is often recommended for about 5 percent of the production piles. 11 POST-FIELD FOUNDATION ASSESSMENT Case Pile Wave Equation Analyses Program (CAPWAP) analyses were performed for each of the 16 tested piles. The CAPWAP analyses, provided an estimate of ultimate pile capacity, the distribution of soil resistance along the length of the pile, and the various soil parameters (such as damping and quake values) required for a revised Wave Equation Analyses. The hammer blow that was selected for analysis was either from near the design penetration depth or near the design capacity based on field PDA data and with representative hammer transferred energy. A plot of PDA versus CAPWAP capacities for both EOID and BOR is shown in Figure 10.
The CAPWAP-generated damping and quake values for each test pile are presented in Figure 11. These data indicate that the damping and quake values vary across the site, with higher values of shaft damping occurring at Pier 5 and between Piers 9 and 12. It has been shown that Smith damping constants slightly decrease for both skin and toe as the soil grain size increases (Likins et al., 1996). Therefore, the variation exhibited across the site might be explained by gradation analyses and index test data which indicate the presence of more fine-grained cohesive soils near Pier 5 and between Piers 9 and 12. For this project site, the quakes and damping values were generally higher than those recommended in the GRLWEAP manual. Consistent with previous experience, it was observed that damping constants during restrike are generally higher than those during the end-of-driving. Toe quakes during restrikes were lower than those at the end-of-driving.
Figure 10: PDA versus CAPWAP Capacities for EOID and BOR. For restrike the earliest representative blow of a warmed-up hammer was selected for CAPWAP. The CAPWAP BOR capacities are also shown on Figure 4. The assessed BOR pile capacities by the CAPWAP are in general agreement with the computed static pile capacities.
Figure 11: Damping Constants (Smith) and Quakes derived from CAPWAP analyses
12 PRODUCTION PILE DRIVING CRITERIA Based on the results of dynamic pile testing at each pier, additional wave equation analyses were performed to develop pile driving and restrike criteria for production piles at each pier. Production pile driving criteria for typical piers (Pier 2 and Pier 11) are presented in Figure 12. The pile driving equipment used by the contractor along with the soil damping and quakes obtained from CAPWAP analyses and measured transferred energies by PDA were used in the wave equation analyses model. These criteria when compared with driving criteria obtained from the pre-field wave equation analyses are less stringent at some locations (Pier 2), and more stringent at others (Pier 11).
underlain by stiff clay. Given the variable nature of setup behavior of these soils, recommending a single value is not prudent. Site specific testing is therefore, essential to determining setup. Plugging did not appear to occur during pile driving for test piles at Pier 2 and 7. It may have occurred during restrike at Piers 11 and 12 as was predicted by static computations and evidenced by increase in BOR capacity. Performance of the hammer was also evaluated. It was found that sometimes the hammer performed erratically and needed some adjustment. Hammer performance can change significantly with time. To ascertain hammer performance during production pile driving and to assure the validity of driving criteria, additional PDA testing should be recommended for the production piles. Production pile driving criteria established from the test pile program resulted in reduced uncertainty and a more reliable foundation. ACKNOWLEDGEMENTS
Figure 12: Pre- and Post- Field Testing Criteria for Piers 2 and 11.
The author wishes to extend special thanks to: Dr. George Munfakh of PB Americas, Inc. for encouraging this work, and especially to Mr. Raymond Castelli and Mr. John Wisniewski also of PB Americas, Inc. for guidance during the field program. Opinions expressed in this paper are solely of the author, and are not necessarily consistent with the policy or opinions of PB Americas, Inc.
13 CONCLUSIONS
REFERENCES
Typical results of a pile installation monitoring and assessment program implemented during the construction of a new bridge in Baltimore are presented. The approach taken to assess the pile foundations, involved 1) preliminary wave equation analyses, using site specific soil information and hammer data, 2) pile driving measurements, 3) post-field CAPWAP analyses, and 4) revised wave equation analyses based on field data and CAPWAP analyses. The static pile capacities predicted using the conventional procedures compared favorably with the field measured EOID capacities and were less than BOR capacities. The field measured blow counts compared favorably with those predicted using wave equation analyses. Set-up factors for the Cretaceous soils at the site in Baltimore were variable and ranged between 1.0 and 1.4. The higher values of setup occurred at locations
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