Pile Load Test and Implementation of LRFD Specifications ...docs.trb.org/prp/11-2891.pdf · yoon,...

Yoon, Tsai and Melton 1

PILE LOAD TEST AND IMPLEMENTATION OF LRFD SPECIFICATIONS: A CASE STUDY OF CAMINADA BAY BRIDGE PROJECT IN LOUISIANA

Sungmin Yoon, Ph.D., P.E. Senior Geotechnical Engineer

Louisiana Department of Transportation and Development 1201 Capitol Access Road Baton Rouge, LA 70804

Phone: 225-379-1935 E-mail: [email protected]

(Corresponding Author)

Ching Tsai, Ph.D., P.E. Senior Geotechnical Specialist



And

James Matt Melton



Text =5,037 words Figures = 10 × 250 = 2,500 words Tables = 3 × 250 = 750 words Total = 8,287 words

TRB 2011 Annual Meeting Paper revised from original submittal.


PILE LOAD TEST AND IMPLEMENTATION OF LRFD SPECIFICATIONS: A CASE STUDY OF CAMINADA BAY BRIDGE PROJECT IN LOUISIANA

ABSTRACT

An instrumented pile load test was performed as part of the foundation design for the Caminada Bay Bridge project in south Louisiana. As part of the load test program, both static and dynamic load tests were performed. The load-transfer curve of test pile was obtained from the strain measurements using the sister-bar strain gages at six locations along the pile shaft. The test pile resistance was determined using the Tomlinson method for cohesive soils and the Nordlund method for cohesionless soils. The dynamic and static load testing results indicated that the test pile did not achieve the desired design resistance. The static analysis model was calibrated based on the observations of pile load testing program. The design pile length was revised to gain benefit from the shallower scour depths for the revised pile design. Due to the low resistance of the test pile, the engineer determined to use dynamic testing on the production piles to ensure adequate resistances. Taking advantage of the static load test of the instrumented test pile, instead of simply using the smaller resistance factor from dynamic tests, the engineer combined the results of the two test methods and used a combination of resistance factors from both the static and dynamic load tests. This paper presents the evaluation of load test results and the rationale used for the resistance factor selection.



INTRODUCTION

Instead of the allowable stress design (ASD) method for foundation design which applies one factor of safety (FS) to account for all uncertainties including the applied loads, site variability, and modeling errors, AASHTO (American Association of State Highway and Transportation Officials) adopted Load and Resistance Factors Design (LRFD) for all federally funded bridges designed after October 2007. The values of safety factor for the ASD are empirically determined and do not accurately reflect the true risk levels. The LRFD method applies risk analysis to account for uncertainties for individual components such as variability of different loads, shaft and tip resistances of a deep foundation. The intent of this design method is to achieve similar reliability among various load components by using various load factors for different type of loads such as dead load, live load, and wind load because of inherited uncertainties in determining the loads. Load factors are used to account for the load effect variability. Similarly, the individual components of the resistances such as shaft resistance and tip bearing for a deep foundation also contain different risk levels due to various reasons such as soil behavior differences for different modes of failure, model simplifications, variation of soil conditions, etc. Resistance factors are used to account for the uncertainty of these factors. The load factor design method has been used for structural designs for more than three decades. Due to higher uncertainties associated with interpretation of soil properties and highly heterogeneous nature of soils, the determination of resistance factors is much more difficult. The geotechnical community falls behind the structural designers in adopting the LRFD concept. To expedite the transition to LRFD, AASHTO provided resistance factors calibrated using load test databases collected throughout the US. The resistance factors obtained from the national database tend to be smaller than those obtained from the traditional practices due to greater variability resulting from geological differences. Realizing this deficiency, AASHTO (1) recommends that states use risk analyses to produce resistance factors from local databases. Many researchers, therefore, have concentrated their efforts in calibrating the resistance factors using different sets of databases (2, 3, 4, 5, 6, 7). However, there are very few publications dealing with implementation issues. This paper presents a case study on the implementation of the LRFD method for the driven pile foundation design.

OBJECTIVES AND SCOPE

The 2007 AASHTO LRFD Bridge Design Specifications (8) specifies a resistance factor of 0.70 and 0.65 for driven piles installed in medium variability soils using a static load test and a dynamic load test, respectively. However, AASHTO does not offer guidance regarding the resistance factors for a pile foundation installed to a different depth from that of the test pile. This paper showcases the application of the LRFD specifications for driven pile design using a combination of resistance factors based on the interpretation of a static and a dynamic load tests. In order to achieve this objective, the shaft resistance distribution and the tip resistance of the static load test were measured using embedded sister bar strain gages and the dynamic test evaluated using CAPWAP (Case Pile Wave Analysis Program) analysis. The final production piles were designed using a combination of the static load test results and dynamic load tests performed on the production piles.



PROJECT DESCRIPTION

The existing Caminada Bay Bridge is located on the southern tip of the Louisiana Highway 1 (LA 1) which connects the mainland to the Grand Isle, a barrier island, near the mouth of the Mississippi River in Louisiana. The existing bridge has been deteriorating due to the severe corrosive environment and it suffered some damages from the recent hurricanes including the hurricane Rita. This bridge is the only access road supplying Grand Isle, a major fishing and tourist attraction. It also serves as the hurricane evacuation route for the Grand Isle of about 5,000 residents. Currently, Grand Isle is used as the major staging area for the clean up effort of the Deepwater Horizon oil spill disaster. The Louisiana Department of Transportation and Development (LADOTD) decided to construct a new four-lane bridge with two 8-foot shoulders on each side to replacing the aging old bridge. The new bridge will be elevated 48 feet (14.6 m) from the mean sea level, allowing boats to pass through. The total length of the bridge is about 4000 feet (1219 m). Figure 1 shows the location of the site and the pile driving activity at the Grand Isle. The elevated bridge results in more than 50 feet of free standing height for the bent piles. Due to bulkling concern, it was determined that 36 inch (0.91 m) precast-prestressed-concrete (PPC) piles were to be used for this project. To reduce the weight and improve handling of the piles, a 22.5-inch (571 mm) diameter void was included.

(a) (b)

Figure 1 (a) Project Location and (b) Pile Driving at Site.

Gulf of Mexico

Caminada Bay Bridge

Grand Isle



GEOLOGY AND SUBSURFACE CONDITIONS

The site is underlain by Chenier and coastal marine deposits of mostly fine sands with interbedded clays. The upper few feet of the soils are influenced by tidal degradation and deposition. Large scour depths of up to 40 feet were recorded at some localized areas from the storm surge during hurricane Rita in 2005. The Chenier is a beach that have become isolated from the sea by strips of generally freshwater marshes, mainly found around south Louisiana. Field exploration programs consisting of standard penetration tests (SPT) in the cohesionless soils and undisturbed Shelby tube sampling were performed in the cohesive soils to characterize the subsurface soil conditions. The standard penetration tests were performed using a manual donut hammer. The laboratory testing program included moisture contents, soil unit weights, Atterberg limits, and unconsolidated undrained (UU) triaxial tests. The soil stratification at the test pile location from the soil boring and results of laboratory tests are depicted in Figure 2. The average mudline elevation at the test pile location is approximately -4.5 feet (-1.4 m) North America Vertical Datum of 1988 (NAVD 88) and the average water elevation is about 0.5 feet (0.15 m). The soil conditions primarily consist of loose to dense fine sands down to an elevation of -60 feet (-18.3 m). Underlying the sands is a layer of soft to medium stiff clays to the depth beyond engineering significance of this bridge construction. The clays are normally consolidated to slightly overconsolidated.

Note: 1ksf=47.9 kPa.

Figure 2 Profile of Soil Classification and Properties.



PILE RESISTANCE CALCULATION

The static axial resistance of driven pile (Qu) consists of the end-bearing resistance (Qb) and the shaft resistance (Qs). The driven pile resistance is calculated using the following equation:

∑=

+=+=n

1isiibbsbu Af.AqQQQ (1)

where qb is the unit tip bearing resistance, Ab is the cross-section area of the pile tip, fi is the average unit shaft resistance of the soil layer i, Asi is the pile shaft area interfacing with layer i, and n is the number of soil layers along the pile shaft. There are many different methods in determining the shaft and tip resistances. The methods described in the following were used to evaluate the pile resistance for this project.

Shaft Resistance α -Tomlinson method (Cohesive Soils)

The α-method is based on total stress analysis for cohesive soils. Τhe pile shaft resistance per unit area of the pile is as follows:

f = α Su (2) where α is an empirical coefficient which converts the cohesion of the surrounding soil to adhesion and Su is the undrained shear strength. The α value typically varies with the strength of the soil. For calculating the test pile resistance for this project, the α values suggested by Tomlinson (9) are used. The shaft resistance (Qs) is calculated as follows:

∫=L

0ds dzC fQ (3)

where L is length of pile in contact with soil and Cd is effective perimeter of pile. For the Caminada Bay Bridge, Cd is 12 feet (3.7 m).

Nordlund method (Cohesionless Soils)

The pile shaft resistance (Qs) is as follows:

∫=L

dDf dzCPCK0

s )sin( Q δδ (4)



qN ′

where Kδ is a coefficient of lateral stress, DP is effective overburden pressure, δ is pile-soil friction angle, Cd is effective pile perimeter, and Cf is a correction factor. In this study the values proposed by Nordlund (10, 11) are used.

Tip Resistance

The tip resistance (Qb) for piles installed in the cohesive soils is calculated as follows:

Qb= Ab Su Nc (5)

where Ab is area of the tip of the pile and Nc is the bearing capacity factor. The magnitude of Nc depends on the depth of pile embedment. For this project, a value of 9 is used.

In cohesionless soils, the pile tip resistance (Qb) can be calculated using the following formula:

qb NqA ′= αbQ (6)

where q is the effective vertical stress at tip level, α is a dimensionless correction factor, and is a bearing resistance factor varying with φ. The values proposed by Thurman (12) are used

in this study.

Test Pile Design Consideration

Thirty-six inch (914 mm) square PPC (precast, prestressed concrete) piles with a 22.5-inch (571 mm) diameter void are used for this project. The factored load of the test pile is 420 kips (1868 kN). This pile represents a design reach of 13 bents in the initial design. The deepest scour elevation within this design reach is -25 feet (-7.6 m) and was used for all pile bents within this design reach even though the scour depths are much shallower for other bents.

In addition to the resistance requirement for the pile penetration design, the Louisiana Department of Transportation and Development Bridge Design Manual (13) specifies that a minimum pile penetration of 25 feet (7.6 m) below the estimated scour elevation. To ensure adequate bearing resistance, two pile diameters below the tip should be similar or stronger than the material at the pile tip. Otherwise, the tip bearing should be reduced.

Figure 3 (a) illustrates a static pile resistance analysis that was conducted using the software DRIVEN ver. 1.2. (14), which uses both the Tomlinson and Nordlund methods for pile resistance calculations. To take the advantage of higher resistance factor, the designer determined that a load test be conducted for the project. Based on AASHTO 2007 LRFD bridge design specifications, 4th Edition (8), a resistance factor of 0.7 was selected. Considering the pile resistance, minimum penetration below scour elevation, and sand thickness, the designer determined that the test pile be tipped at the elevation of -60 feet (-18.3 m). A casing was to be installed to the scour elevation to simulate the scour condition.



The target pile resistance for the test pile was calculated by taking into account the overburden effect. Figure 3 (b) shows the calculated static pile resistance assuming that a casing tipped at an elevation of -25 feet (-7.6 m) is used. The calculated target pile resistance for the test pile is 1129 kips (5022 kN). Prior to the load test, the designer had expected that a slightly higher pile resistance (1219 kips or 5422 kN) could be achieved. The variability of the calculated pile resistance was evaluated using the graphic three-sigma rule (15). One standard deviaton of the calculated resistance is about 336 kips (1495 kN). The calculated 95 % confidence interval for the calculated resistance is thus between 890 kips (3959 kN) and 1548 kips (6886 kN).

(a) (b)

Figure 3 Static Pile Resistance Analyses: (a) Analysis with Total Scour and (b) Analysis with Casing of 20.5 ft (6.2 m).

PILE LOAD TEST PROGRAM

Instrumentation

In order to better understand the distribution of pile shaft resistance, it was decided that sister-bar strain gages be installed along the test pile. Twelve resistance type sister-bar strain gages were installed in pairs at 6 locations as shown in Figure 4. Pile Installation and Pile Load Tests

The test piles were driven using a Pileco D100-13 single-acting (open ended) diesel hammer with a ram weight of 22.0 kips (97.9 kN) and a maximum rated energy of 246.8 kips-ft (334.7 kN-m) at full stroke. Preliminary Wave Equation Analyses (WEAP) results indicate that the Pileco D100-13 hammer would be capable of installing the 36 inch (914 mm) square PPC pile to a wide range of resistances (approximately 190 to 2000 kips (845 to 8896 kN)). The hammer



performance was monitored using a pile driving analyzer (PDA) during pile installation. The test pile was driven to the design tip elevation of -60 ft (-18.3 m) successfully.

A static compression load test (quick load procedure) was performed on the test pile 27 days after initial pile driving. A 1080 kips (4804 kN) load cell was used for the load test and two dial gages attached to the reference beams were used to measure the pile head movement. Dynamic pile tests using a Pile Driving Analyzer (PDA) were performed to monitor the initial pile driving, 7-day restrike (7-day BOR), and after static load test restrike (ASLT). CAPWAP analyses were performed for the end of initial drive (EOD) and the two restrikes. (a) (b)

Figure 4 (a) Layout of Test Pile Instrumentation and (b) Installed Sister-bar Strain

Gages. RESULTS

Driving Resistance

The driving resistances versus elevations were recorded using a PDA during the test pile installation. The results are presented in Figure 5 along with the maximum Case pile resistances (RMX). A Case damping constant (JC) of 0.5 was used during pile driving due to the existence of predominant fine sandy soils. The maximum pile resistance estimation in the sand layers



appears reasonable. However the base resistance at the end of driving may be overestimated because a Case damping constant of 0.5 may be too low for the clay soil at the pile tip.

The resistance profiles clearly indicate that the test pile penetrated the dense sand layer located at an elevation of about -50 feet (-15.2 m) and tipped at a clay layer at about elevation -60 feet (-18.3 m). The soil stratifications inferred from the driving resistance match closely to those observed from the soil boring log shown in Figure 2.

Figure 5 Pile Driving Resistance with Elevation. Dynamic Load Test Results

CAPWAP data analyses performed with PDA data obtained under representative hammer blows from each test set (EOD, 7 days BOR and ASLT) are summarized in Table 1. Figure 6 shows the wave matching analyses and the estimated load-settlement curves using CAPWAP. As can be observed in Table 1 and Figure 6, the base resistance remained virtually the same even after 27 days. The shaft resistance, however, increased from 251 kips (1117 kN) to 406 kips (1806 kN), nearly doubled. The increase in the resistance between 7 and 27 days was only about 46 kips (205 kN) for the shaft and 30 kips (133 kN) overall, relatively small amounts, suggesting a logarithmic time relationship similar to the model proposed by Skov and Denver (16). The logarithmic time capacity plot for this pile is shown on Figure 7 and is consistent with the observations made by many researchers (17) including several projects in Louisiana (18, 19).



(a)

(b)

(c) Note: 1 kip=4.45 kN and 1 inch=25.4 mm

Figure 6 Wave Matching Analyses and Derived Load Displacement Curve Using CAPWAP: (a) EOD, (b) 7 days BOR and (c) ASLT.



Table 1 Summary of the Characteristic of the Investigated Piles

Dynamic Load Tests Total Shaft Base Kips kN kips kN kips kN

End of Initial Drive (EOD) 450 2002 251 1117 199 885 7 Days Restrike (7 days BOR) 570 2535 360 1601 210 934 Restrike After Static Load Test (ASLT) 600 2669 406 1806 194 863

Figure 7 Normalized Total Pile Resistance versus Time Ratio

Static Load Test Result

The applied axial load and pile head displacement is presented in Figure 8. The pile sustained an axial load of 558 kips (2482 kN) with a pile head displacement of only about 0.5 inch (12.7 mm). The pile then plunged with a slight increase in the applied load. The ultimate pile resistance was interpreted to be 540 kips (2402 kN) using the modified Davisson method proposed by Kyfor et al. (20) for pile size greater than 24 inches (610 mm). The interpreted resistance is about 2.1 standard deviations lower than the calculated mean resistance, which demonstrates a significant deficency. The static load test result showed a close agreement with the dynamic load test resistance of 600 kips (2669 kN) (ASLT). The pile resistance difference determined from the static and dynamic tests is only about 10% (60 kips or 267 kN) well within the accuracy of the test methods. Note that the Davisson’s criteria produced a smaller ultimate resistance from the plunging load and that the dynamic testing is supposed to produce ultimate resistance close to the



plunging load. Therefore, the proper comparison of the dynamically determined resistance should be the plunging load. In this case, the difference is only 42 kips (187 kN), about 7%.

Figure 8 Load-Settlement Curve.

Shaft and Base Resistance Interpretation using Strain Gages

Figure 9 shows the load transfer plots corresponding to each applied load increment from the static load test. The top and bottom strain gages were located 5 feet (1.5 m) below the casing elevation and 2.3 feet (0.7 m) from the pile bottom, respectively. In order to properly determine the skin friction distribution, it is assumed that the unit skin frictions do not change above the top gage and below the tip gage. The complete load transfer curve was, therefore, obtained by extrapolating the resistance to the pile ends. The residual stress was determined to be insignificant due to the soil condition and relatively low base resistance. As shown in Figure 9, about 82% of the total pile resistance is from the shaft resistance. The interpreted shaft and base resistances at the plunging load are 460 kips (2482 kN) and 98 kips (436 kN), respectively. The interpreted base resistance is lower than the derived base resistance using the wave matching method while the shaft resistance is 15 % greater than the dynamically determined value. There are two possible explanations for the differences. First, the dynamically determined tip resistance value may be too high. This is evident by the statically calculated value of about 90 kips (400 kN) based on the clay soils at the tip. The other explanation is that the measured tip resistance is for the plunging condition where strain softening might have reduced the tip resistance from the peak resistance.



Figure 9 Load Transfer Distribution for Each Loading Increments.

PILE RESISTANCE ANALYSIS AND LRFD IMPLEMENTATION

Measured and Predicted Resistances

Table 2 shows the predicted pile resistance using static calculation methods and the results from the static and dynamic load tests. As can be seen in the table, the static calculation significantly overpredicted the pile resistance. To achieve the desired 1129-kip (5022-kN) factored resistance, additional pile length will be needed for the production pile or a new test pile program be conducted. However, considering the cost and schedule, an additional pile load test is an unlikely solution. To investigate the causes of the large variation between the predicted value and load test result, the authors concluded the following possibilities. First, the SPT N values obtained using a manual donut hammer resulted in higher blow counts than actual soil densities. The average energy ratio (Er) of a rope and pulley donut hammer is reported to be 45 about 50% of the energy ratio of a safety automatic hammer (Er of 80-100) (21). If the blowcounts were to corrected to the standard N60 value, the reported N values should be reduced by roughly 20 percent. Thus, a much smaller friction angle for the pile resistance calculations should have been used. Second, the Nordlund method might overestimate pile resistance for large diameter piles because the method was derived from the observations on piles with size ranging from 10 inch (250 mm) to 20 inch (500 mm) only (14). Extrapolating pile size to almost twice of the

Load Applied to Pile Head



originally studied piles might result in the large errors observed in this project. As such, it was concluded that the static calculation method should not be used for production pile order length determination without adjustment to the soil parameters.

Table 2 Evaluation of Test Pile Resistance

Pile Resistance Total Shaft Base

Deviation from Required Total Resistance of 1129

kips (5022 kN) kips kN Kips kN Kips kN Kips kN

Static Analysis (Tomlinson method

and Nordlund method) 1219 5422 1189 5289 30 133 +90 +400

Dynamic Test using CAPWAP 600 2669 406 1806 194 863 -529 -2353

Static Load Test Combined with Strain

Gage Monitoring (Plunging Load)

558 2482 460 2046 98 436 -571 -2540

Static Load Test (Davisson Criteria) 540 2402 - - - - -589 -2620

Implementation of LRFD Specifications

Based on the observations on the test pile and analyses of the results, the DRIVEN model was calibrated to match the interpreted shaft resistance distribution based on strain measurements. Because the test pile was driven only through in the fine sand layers to -60 ft (-18.3 m) elevation, the DRIVEN model calibration was made on the layers that the test pile was driven in as shown in Table 3. The reduced frictional angles of the calibrated values appear to be consistent with the inferred low hammer energy of the SPT hammer. Figure 10 illustrates the shaft resistance profile from the static and dynamic load tests and static analyses from both the calibrated and original model. As can be seen in the plot the calibrated DRIVEN model matches well with the static load test result.



Table 3 Calibration of Soil Strength Parameters

Layer Elevation Interpreted Friction Angle, φ (o)

(ft) (m) Initial φ Calibrated φ 1 -4.5 to -14 -1.4 to -4.3 28.1 28.0 2 -14 to -19 -4.3 to -5.8 29.5 25.0 3 -19 to -25 -5.8 to -7.6 31.0 28.0 4 -25 to -37 -7.6 to -11.3 33.9 26.0 5 -37 to -46 -11.3 to -14.0 35.0 26.0 6 -46 to -60 -14.0 to -18.3 37.7 29.0

0 890 1779 2669 3559 4448 5338 6228

‐19.8

‐18.3

‐16.8

‐15.2

‐13.7

‐12.2

‐10.7

‐9.1

‐7.6

‐6.1

‐65

‐60

‐55

‐50

‐45

‐40

‐35

‐30

‐25

‐20

0 200 400 600 800 1000 1200 1400

Resistance (KN)

Elevation (m

)

Dist. from

Bottom (ft)

Resistance (kips)

Shaft Resistance (Strain Measurement)

Calibrated DRIVEN

CAPWAP Resistance

Initial DRIVEN

Figure 10 Shaft Resistance Profile from Static and Dynamic Load Test. To improve the accuracy of determining pile order lengths, the geotechnical engineer decided to reduce the number of pile bents for this design reach in addition to the DRIVEN model calibration to take the advantage of shallower scour depths for the smaller design reach. The long term scour and pier scour depths used for revised final design are 4.0 ft (1.2 m) and 9.7 ft (3.0 m), respectively. The contraction scour was not considered in the final design due to the small scour depth of this component. As expected, the gain in overburden from the shallower scour shows a significant impact on a pile static resistance in cohesionless soils.



The basic LRFD principle can be expressed in an equation form as follow,

∑∑ ≥ iin QR γηφ (7)

where φ is the resistance factor, Rn is the nominal resistance, η is a load modifier to account for effects of ductility, redundancy and operational importance, Qi is the load effect, and γi is a load factor. For this project, the factored load effect is 420 kips (1868 kN). The shaft resistance above elevation -60 ft (-18.3 m) was verified using the static load test. AASHTO (8) specifies a resistance factor of 0.7 for the static load test. However the pile resistance below the elevation of -60 ft (-18.3 m) was not tested. Considering the uncertainty of the DRIVEN model without load test calibration and the close agreement between the dynamic and static tests, the authors decided to perform dynamic testing on the production piles to control the pile penetration. AASHTO specifies a resistance of 0.65 for piles installed using dynamic testing. It was decided that a resistance of 0.7 be used for the pile resistance above the elevation of -60 feet (-18.3 m) and a resistance factor of 0.65 be used for the pile resistance below that elevation. The final pile tip elevation will be determined based on the dynamic test result of production piles using the above combination of the resistance factors. Based on the laboratory test results of the clay soils below the -60 feet (-18.3 m) elevation, the tip of the production pile is estimated to be at about -78 feet (-23.8 m), about 18 feet (5.5 m) longer than the test pile.

Predicted Shaft

Resistance above -60 ft Elevation

×

+

Predicted Shaft and Base

Resistances below -60 ft Elevation

×

≥

Factored Load Effect (420 kips/ 1868 kN)

(8)φ(0.7) φ (0.65)

As a comparison, a calculation was made to determine the pile length required if no dynamic testing is used. AASHTO (8) specifies a resistance factor of 0.35 for redundant piles when using Tomlinson’s a method in clay soils. The required pile tip elevation became -95 feet (29.0 m), which is 17 feet (5.2 m) deeper than the one determined based on the dynamic load test. The pile length requirement for this pile tip elevation exceeded the contractor’s crane capacity even when the pile is spliced. The cost associated with additional mobilization, crane rental, and pile casting and installation became unacceptable. FINDINGS AND SUMMARY

This paper presents a driven pile design process based on the LRFD concept. The pile resistance was initially predicted using the Tomlinson method for cohesive soils and the Nordlund method for cohesionless soils and was found to be unreliable. The total pile resistances and load-transfer distribution interpreted from the dynamic and static testing methods are in close agreement. Due to the overestimation of the resistance during the test pile program, the test pile does not



represent the final production pile lengths. A method of using combinations of resistance factors is proposed. The following discussions summarize the findings from this case study.

1. Bridge scour has a great impact on the pile foundation design, especially in cohesionless soils. A 7-foot (2.1 m) reduction in scour depth for the smaller design pile length allows a nearly 20 percent gain in the usable pile resistance.

2. The test pile was embedded almost entirely within the sand layer with only 2 feet embedment into the clay soil at the pile tip. Although the most of skin frictional resistance was provided by the sandy soil, the skin frictional resistance nearly doubled after 27-day setup period. The results show that the majority of the setup effect was mainly developed in early period after pile driving. However, the pile continued to gain resistance upto 27 days, the final load test performed.

3. The driven model overpredicted the pile resistance most likely due to the low efficient SPT hammer. The use of properly calibrated SPT hammer is essential in predicting pile resistances. Should a load test have not conducted for this project, the foundation would have been severely under designed resulting in an unsafe condition.

4. The total pile resistance determined using dynamic load test was within 10% of the static load test result. The static and dynamic load tests, when properly analyzed, produce comparable results.

5. When the static load test fails to achieve the desired resistance, using a combination of the test results and statically calculated resistance or dynamic test can compensate for the difference. A caution should be taken when applying this implementation in order to revise a design pile length. This implementation should be exercised only when the geotechnical engineers have intimate knowledge of the soil conditions below the elevation tested using static load test and the load transfer distribution of test pile.

6. The use of dynamic test for pile length control can produce significant savings over the static calculations on the pile lengths based on the AASHTO’s resistance factors.

7. The geotechnical engineer should be aware of the potential shortfall of the Norlund’s method for large diameter piles.

ACKNOWLEDGEMENTS

The authors gratefully acknowledge the support of Steve Meunier, Administrator of the Pavement and Geotechnical Section, LA DOTD. The assistance of Dane Drew and Chris Rome in preparing the graphs is acknowledged.



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