3.0 Literature Review on Different Design Methodologies

3.1. Introduction

In the first part of this chapter, the main focus is to give the reader a brief idea on skin

friction, PDA® testing, reliability of PDA® based on globally accepted researchers and

finally, the reliability of PDA® results obtained from Ceylinco Celestial Project for the

purpose of this case study. The second part of the chapter will be to introduce the design

methods used to evaluate pile capacities in Chapter 4.

3.2 Skin Friction

As the pile is loaded axially and forced to move downwards, the first type of

resistance it has to undergo is the skin friction (Bowles, 1996 and Tomlinson, 1994). The

skin friction is activated under very small displacement (Tomlinson, 1994) and its magnitude

depends on the strength properties of the surrounding soil, method of the installation of the

pile. and the properties of the pile surface etc. (Poulos and Davis, 1986).

Skin friction is mobilized in both cohesive and cohesionless soils and can be

calculated by various methods (Bowle, 1996; Tomlinson, 1994 and Poulos and Davis, 1986).

Skin friction in cohesive soil is not discussed in this dissertation as the site of the case study

considered does not contain clayey soil in the subsurface. This research is mainly focused on

estimation of the ultimate skin friction resistance developed in cohesionless soil layers on

bored and cast in-situ concrete piles in Sri Lanka.

In order to estimate the skin friction of bored and cast in-situ concrete piles, certain

data must be obtained, such as unit weights, soil strength properties and the lateral soil

pressure applied on the pile by the soil along the pile shaft. As mentioned before, the method

of the installation of the pile has a major impact on the skin friction developed on piles as the

strength properties of the soil surrounding the pile and the lateral pressure on the pile shaft

from the surrounding soil depend on the method of the pile installation (Poulos and Davis,

1986).

Due to the importance of the contribution of the skin friction on the total capacity of

piles, many research studies have been conducted throughout the world and research articles

and books are published on the estimation of skin friction on piles. Based on these research

studies, there are large number of empirical equations and design criteria published by many

professionals on this subject. When compared with other countries very few research studies

have been carried out in Sri Lanka related to estimation of carrying capacity of bored and

cast in-situ concrete piles. In most cases design methodologies and the parameters are

directly obtained from foreign literature to be used for design of piles in Sri Lanka. As

previously mentioned, the pile installation methodology plays a vital role in the development

of the skin friction. Therefore, in this research, it is aimed at studying the development of

skin friction on bored and cast in-situ concrete piles in sandy soil and identifying the best

possible way to predict skin friction to the acceptable accuracy.

In this endeavor of establishing the possible method of estimation of the skin friction

developed on bored and cast in-situ piles in Sri Lanka, it is essential to compare estimated

skin friction resistance from different methods with the measured skin friction resistance

actually mobilized on piles. There are various methods adopted by researches in the other

countries to measure the mobilized skin friction on piles during pile load tests. Among them,

instrumented pile load testing is very common. Furthermore, advanced pile load tests such as

Osterburg cell load testing gives high accurate estimating of the mobilized skin friction on

piles. In these methods, the skin friction mobilized on the pile is directly measured during

loading of the pile and the measured skin friction is highly accurate. In spite of the high

accuracy levels obtained from such testing methods, these methods are still not in used in Sri

Lanka. Because of that, comparatively less accurate high strain dynamic load testing of piles

using the Pile Driving Analyzer (PDA®) used in this research study. The accuracy ofthe skin

friction determined from the P D A® testing is described in subsequent sections of this thesis

together with the method utilized to check the accuracy level of the determined skin friction

resistance from the PDA®testing.

3.3 High Strain Dynamic Testing of Bored Piles

High strain dynamic testing of bored piles was introduced to Sri Lanka in late 1990's

and dynamic load testing of piles using Pile Driving Analyzer (PDA®) is now very widely

used in Sri Lanka. The skin frictional and end bearing components of the developed

resistance could be determined by the dynamic testing of piles using P DA®.

In the high strain dynamic pile load testing using a Pile Driving Analyzer (PDA®), a

pile instrumented at the top with two accelerometers and strain gauges, is impacted with a

heavy hammer so that a high strain stress wave propagates along the pile shaft. During the

blow, the acceleration and the strain in the pile at about 1.5 x pile diameter are recorded

using the accelerometer and the strain gauges, attached directly on to the diametrically

opposite sides of the pile as shown in Figure 3.1.

(a) (b)

Figure 3.1 (a) Strain gauges fixed to sides of a pile and (b) close up view oft he strain gauges

The measured strain is converted to force using cross sectional area and elastic

modulus of the pile and acceleration is integrated to obtain the velocity of the pile top.

Therefore, the velocity and the force developed at the pile top during the hammer blow can

be obtained. These measured velocity and force history at the pile top during a hammer blow

are used to simulate the behavior of the pile during static loading. It should be noted here that

the pile is dynamically loaded during the dynamic load testing to obtain the response of the

pile under static loading. For detailed description regarding the dynamic load testing of piles,

the reader is referred to Thilakasiri et. al. (2006)

Mainly two methods are used in the analysis of the velocity and the force

measurements at the pile top during a hammer blow to predict the static capacity of the pile:

(i) using the wave propagation theory and assuming a uniform elastic pile and ideal elastic­

plastic soil behavior and; (ii) using combined wave equation soil model and a continues pile

model to iteratively determine the unknown soil parameters by matching the field recorded

velocity and force measurements at the pile top. The analysis mentioned in the first method

could be done easily in the field and an approximate static carrying capacity of the pile and

any defect in the pile shaft could be obtained for assumed soil parameters. The measured

pile top velocity and force records can be downloaded to a computer and a more rigorous

matching process could be used to obtain skin friction distribution along the pile shaft,

mobilized capacity at the pile toe, simulated pile static lead settlement curve, cross sectional

variations along the pile shaft and, soil damping and stiffness parameters.

Variations of the acceleration and strain, during the application of the hammer blow,

are obtained in the field using the PDA and the acquired data is processed to produce velocity

vs. time and force vs. time records at the pile top during hammer blow. Subsequently, the

data is transferred to a personal computer and a rigorous numerical analysis (modeling)

procedure called the Case Pile Wave Analysis Program (CAPWAP), which is formulated

based on the wave equation method, is used to analyze date to obtain the static Load vs.

Settlement curve and other information related to the pile.

In the CAPW AP analysis. the pile is first divided into number of elements

considering the wave speed in the pile. The simulation technique, used in the CAPW AP

method, is referred to as the wave equation method first introduced by Smith (1961). The

detailed discussion on the wave equation method is beyond the scope of this study and the

readers are referred to Thilakasiri et al.(2003) for more details of this numerical procedure.

At the beginning of the analysis. a complete set of wave equation and type soil parameters,

such as damping coefficient of the skin CJsk 111 ), damping coefTicient at the pile end CJend), skin

quake and end quake, are assumed and entered into the computer program. Then, in the

dynamic analysis, the measured velocity is imposed on the top pile element and CAPW AP

calculates the force necessary to induce the imposed velocity using the wave equation

method. Typical measured and estimated force at the top of the pile during a hammer blow is

shown in Figure 3.2.

80 to -- Force measured --- Force computed

40 1

0 --..;.__-- 1 L

Figure 3.2 Measured and computed pile top .force variations obtainedfrom a PDA® test

Cl'SION IIIIi ''''''''1 BLOCK

(!) u c Zl U1

Vl (!)

c::: c .s u ·.: u...

0 [/)

..'!::! 0::

Actual Pile

CAP BLOCK

Soil Dashpots , •I

W(IO)

Smith Pile Model

Figure 3.3 Smith Idealization o.f Pile Soil System

Pile springs

Springs

A parameter called "Match Quality" is used to indicate the closeness of the match

between the measured and the estimated force at the top of the pile. If the two curves are

identical the match quality is zero and will increase with the difference between the measured

and the predicted curves. It is generally found that if the match quality is less than 5, it can

be considered as an acceptable match (PDA® Manual). As long as the match between these

two quantities is unsatisfactory, the process of iteratively changing the parameters of the soil

model will continue.

The CAPW AP simulation is an iterative procedure to match the measured force or

velocity with the same from the predictions from the wave equation method.

Thilakasiri et. al. (2006) investigates the accuracy of the load-settlement behavior

from the CAPW AP simulation and showed that the predictions can be improved by

considering,

1. The applicability of the skin friction from the CAPW AP simulation to the soil

profile at the tip.

2. Measure blow count and the estimated blow count from the CAPW AP simulation.

3. The dynamic soil parameters used in the simulation process.

It has been shown by various researches by comparing the results of static load test

and the dynamic load test done on the same pile, the load settlement behavior predicted by

the dynamic testing method is accurate (Globe et aL 1980, Links et al. and Rausche, 1980)

Few case studies are discussed below in order to show that the results predicted by

PDA® (CAPWAP analysis) are comparable with static load test results, and thus it would be

reasonable to assume that the static friction capacity predicted by the dynamic load testing of

piles is also reasonably accurate.

It has shown by various researchers that the results of dynamic load test do match

with the results obtained from static load test to the acceptable levels. But to achieve this

match there are certain criteria to be followed during testing which are discussed below. So,

it is worth to look at the correlation of CAPW AP analysis results with static load test results.

Although there are many applications for dynamic pile testing, bearing capacity is

being the main one. The ability to predict the static capacity from dynamic pile testing has

resulted in many studies and has been the focus of dynamic pile test on many project sites.

Standard practice requires performing single matching on the data to more accurately

determine capacity from the dynamic tests. (Links et. al. and Rausche, 1980)

Reliable correlations for long term capacity from dynamic tests with static load tests

require simple guidelines. For driven piles, dynamic tests should be performed during a re­

strike after a sufficient wait period to allow soil strength changes to stabilize. Testing of

drilled shafts or cast in-situ piles requires the concrete or grout to achieve a sufficient

strength, which indirectly allows the soil to recover from the drilling process. The driven or

drilled piles must also experience a reasonably net set per blow to mobilize the full capacity.

Since dynamic testing of drilled shafts offers results in a small set per blow, the capacity

mobilized would be less than the respective ultimate capacities.

Based on the original research work at case Western Reserve University under the

direction of Dr. G.G.Globe, the CAPW AP analysis procedure was both developed and

reported (Globe, 1980).

Correlation of Globe 1980 study resists of H piles, timber piles and concrete piles are

given under Figure 3. 4 ( Links et. al. and Rausche, 1980).

I !

usru\.; ' ,,;J~'~·;~ · :~·~ ~w'i[!~!·' ; 1~·-

'\.~(}~-~·~-·- (t 0

............................ l ........................... l'''''"'"'':·············

' It t •

3000

2000 '· I • ~' • .......... . ' . .. ·'. •

1000 ~ ¥¥1' I

o~----~-----4~--~

0 1000 2000 300(

ILT [\Nl

Figure 3.4 CAPWAP (CW) vs. Static Load Test (SLT). Correlation of Goble's 1980 Study.

This was one of the earliest studies done to find out the correlation of the capacities

predicted by CAPW AP with static load test results. Though the data consists of different

types of piles it confirms the idea that the CAPW AP results closely match with the static load

test results. An accurate comparison should be carried out by performing static and dynamic tests

with a smaller time gap between them, as the pile capacity changes with time. Thilakasiri et

al. (2003) carried out a study on the strength of driven piles and showed that a considerable

gain of capacity with the time. But due to the lack of information about the date of

installation or cast. the research results will not show the true representation of the actual

scenario, since the pile capacity generally changes with time. Proper evaluation of capacity is

a time dependent effect. Unfortunately only slightly less than half of the cases contained

information on date of dynamic tests and static testing relative to the installation date.

Inclusion of dates allows computation of the "Time Ratio", defined as the time of the

dynamic test divided by the time of the static test, both relative to the casting date. So, that

time dependant soil strength changes after installations are minimized, a Time ratio of 1.0 is

usually ideal. (Likins et. al. and Rausche , 1980).

So the "Time ratio" is very crucial in the comparison of CAPW AP results with SLT.

Thus date of tests, relative to the date of pile cast, should be included in future reporting of

results.

q ? (' 0 '" .., ...JJ 'v

The following examples of time ratio> 0.25 give a good correlation. A comparison of

the CAPW AP results with SL T load shown in Figure 3. 5 below. Correlation of results is

excellent and rejects the doubts on the precision ofthe CAPWAP calculation, stiffness ofthe

pile, soil system and soil resistance distribution. (Likins et. a!. and Rausche, 1980)

20.000 -··--·-····-··--·"·-··· ·----·--~--.-··

16.000 -+-----1---t--+---

~o.ooo 1 1 .1 f •

!5,000 I "t. I I

0 5.000 10.000 15.000 20.00C

SLT l'kH1

Figure 3.5 CAPWAP vs. SLT time ratio [CWISLT > 0.25]

This shows that to have good correlation between two results, it should have

considerable set in dynamic testing and relatively higher settlement in static testing .

..0.000

30,000

[ 20,tm

~ 10,000

0

---.....-.-i

! ·• i --1

i i j

• I !

• l • • I i __.

•• I .;

• i

,~ • ! ;

0 10.000 20.000 30.000 .co.ooc SLTikHJ

Figure 3. 6 CAPWAP vs. SLT drilled and auger cast piles

Figure 3. 6 presents the result for cast in situ drilled and augured piles. Drilled cast in

situ piles show less correlation compared it with the driven piles. This is due to the fact that

driven pile has more reliable information of both the shape (e.g. Cross sectional area vs. pile

length) and modulus of elasticity. For drilled shafts this forms may vary with the lengths.

hen though the drilled shaft has these disadvantages the results show that CAPW AP and

SLT results are quite closer to each other.

By looking at above results shown from the researches conducted by Garland Links

and Frank Rausche (1980) on correlation of CAPW AP with SLT results following

conclusion can be derived.

To have a good pre calculation of capacity from CAPW AP to match with SL T proper

guidelines on time and net settlement of pile testing has to be followed. But by all the test

results it clearly shows that if the above criterion is satisfied then the values predicted by

CAPW AP shall be accepted inline with SL T results.

Before the use of PDA® results to estimate the skin friction component developed in

bored cast in-situ piles, it is advisable to prove the applicability of PDA® method to estimate

the mobilized skin friction. For this purpose, the methods proposed by Van Weele (1957) and

Chin (1979) are used. Applicability of the method proposed by Van Weele to estimate the

skin frictional capacity from static pile load test results was researched by Thilakasiri (2006).

So it is beyond the scope of this research and here it is assumed that Van Weele (1957)

method estimates the skin friction component to the accuracy required for the purpose of this

research. But a comparison between skin friction calculated from SL T using Van Weele

method and skin friction values obtained from PDA® testing are discussed under following

paragraphs. In addition to those two methods results obtained from defective piles of

Ceylinco Celestial Residencies Project (2003-2009) has been used to build an argument on

the reliability of PDA® results.

3.4 Investigation of the reliability of the P DA® results

3.4.1 Results from defective piles

Tahle 3.1 PDA® resultsfor defective piles. CCRP (2003-2009)

,--~ --

I P D A® before rectification l P D A® after rectification

f--- -Skin Friction ! End Skin End

(MT) Bearing(MT) Friction(MT) Bearing(MT)

1--- -

P049 (1800mm) 611 700 1875 . 715

P114(1800mm) 854 683 1811 842

f----

P123( 1800mm) 548 576 1686 1373

P072( 1800mm) 775 411 1782 1173

P142(1800mm) 548 589 2125 451

f------P149( 1500mm) 597 420 1570 394

f------P074(1800mm) 685 278 1525 375

P087(1800mm) 1185 693 1686 1292

'------

Above table gives the results of skin friction and end bearing obtained for "Defective Piles",

before and after the rectification. The term "Defective Piles" describes the piles of which the

capacities are not mobilized to the tested values.

The defects which were identified by the experts was "soft toe" due to sedimentation of sand

or other loose materials at socketed portion and toe area of the pile. The method used to

identify the suspected piles was a simple technique.

All the piles in above project were tested for PIT®. The Low Strain Pulse Echo method of

testing the integrity of the piles is accepted world wide. The test method has been codified by

professional institutions in many countries and testing standards and codes have been

developed for its proper specification and use. When a cast in-situ pile is struck with a small

hammer, a stress wave is generated that travels down the shaft to the bottom where it is

reflected. When the reflected stress wave returns to the top, a measurable pile top motion

occurs. If the reflected wave occurs at the correct time, and if no other earlier reflection is

observed at pile top, then pile shaft is likely to free from any major defects. Then series of

analysis done with plotted curves, it can be concluded whether a pile is a defective one or

not. Sine the decisions are heavily depend on the interpretation of individuals it requires

some experienced personal in the piling industry to take decisions. Identification of defects

~

was mainly done using the pile integrity testing method briefly described above and using a

point scheme to decide the defectiveness of piles which is described below.

Time taken for boring of a particular pile, difference between actual and theoretical concrete

volumes. time gap between the completion of boring and start of concreting and PIT® test

results are the considerable factors for quality of piles. They were given marks by

considering the importance of them to the quality of piles. For an example PIT® results

giYes the priority and if a pile suspect as a defective one by PIT®, then it is assigned full 10

marks. Similarly other factors also give marks and obtained total marks for individual piles.

A pile with higher marks means the higher probability to becoming a defective pile. Then

those selected piles are tested for P D A®.

If the mobilized capacity of piles is less than the expected value then the pile is suspected as

a defective pile. Post grouting technique using "U" tube method was used to rectify those

piles and piles were re-tested using PDA®. In the above method two cylindrical vertical holes

were bored in the defective piles at predefined spacing as indicated in the Figure 3.8. Core

recovery is closely examined by an expert during the process. When it reaches to the weaker

area of the pile. in this case to the toe, other borehole drilling can be started. When the expert

confirm that both are at same weaker are of pile then high pressure water jet of about 20 bar

applied to one end while tightly closing the other end. After pressure reach it predefined peak

value then close cap of the second end release. Water jet with impurities comes out at a

greater pressure and process continues until water comes out from the pile is becoming clear.

After removing all loose materials, dry air jet of high pressure applies in to the drilled hole to

make it dry. Then grout prepare with ordinary cement is applied in to the drilled hole until it

is completely filled under pressure. Allow 14 days for curing and then that pile can be tested

using P D A® again.

l • l ,.

: ~ t j

l

:II

li

Drilled hole Nol

75mm Diameter

High pressure water in

Concrete pile

Drilled hole No2

u u

Figure 3. 7 Rect(llcation ofpile using "U" tube method

High pressure water out

GL

Concrete pile

Contaminated area

Loose area Filled with Grout

Contaminated area

First column of the above table 3.1 gives the SF before rectification for each pile and the

average is around and this may be the "Ultimate skin friction" generated by the top

soil layers up to rock strata. This considered as the "Ultimate skin friction in sandy soil

region", because the piles have moved down excessively due to the defects in the toe and

socketed region. Reports suggested that generation of SF in this weaker socketed area is very

low and can be considered as zero. Thus the excessive movement should mobilize the

ultimate skin friction in sandy soil region. After grouting this figure has increased by an

average of 1 but there are only three cases out of eight the end bearing figure increased

in large percentage. It suggests that grouting of toe really effected on the SF of the piles

rather than EB. Since it has done only the toe rectification, it is very clear that the increment

of SF is produced by the socketed portion of the piles.

To ensure the above argument, the average values of skin friction up to the rock strata given

by the five design methods discussed in Chapter 4 are considered. Value given by them is

900MT and values produce in the actual case is reasonably close to the predicted values by

the PDA

As the fl.rst conclusion of the argument given above, the "Ultimate skin friction up to rock"

in this particular case excluding the socketed area for 1800mm diameter piles can be taken as

approximately 600 MT.

The average value SF after the rectification is 1600 MT. The individual values are very

similar and close to each other. That means after proper grouting it can assume

comparatively proper pile geometry inside the socketed area. This increases the probability

of producing similar results by same diameter piles. So that the average of this results may

represents the true skin friction capacity of 1800mm diameter piles of this particular site.

As the second conclusion of the argument given above for this case study, the "Ultimate skin

friction of the pile'' including the socketed portion for 1800mm diameter piles can be taken

as approximately 1600 MT.

These two conclusions will be used in following discussions on Van Weele method ( 1959)

and Chin Method ( 1979) in following chapters.

3.4.2 Discussion on the application of Van Weele (1957) method

In this section Van Weele (1957) method and its application to this case study would

discuss to verify the results obtained from PDA tests are relial:5le. The values obtained from

the PDA tests are compared with the results of Van Weele (1957) method.

By traditional load~settlement curve obtained from a static load test, the load

settlement behavior at top of the pile could be obtained. Even though a normal load­

settlement curve doesn't directly give the skin friction and end bearing separately, the shape

of the load settlement curve depends on the relative magnitude of the skin friction and end

bearing and the distribution of the skin friction along the pile shaft. According to the Van

Weele (1957) and Bowles (1986), when pile is loaded initially the load is carried mostly by

skin friction until the shaft slip is sufficient to mobilize the limiting value. When the limiting

skin friction is mobilized, the point load increases nearly linearly until the ultimate end

bearing capacity reached. At the point of the ultimate end bearing, the load settlement curve

becomes vertical indicative of no additional load carried due to further increase in settlement.

Based on the above argument, a typical load-settlement curve, shown in Figure 3. 8 has three

distinct regions as below:

1. Region from point 0 to A - Within which the capacity is mainly from the skin friction plus

small contribution from end bearing.

2. Region from point A to B - Within which the load capacity is the sum of the limiting skin

friction plus the approximately linearly increasing point bearing capacity.

3. Region from B to point where curve become vertical. Often the vertical asymptote is

anticipated and often the test is terminated before a "vertical" branch is established."

(Thilakasiri, 2006).

By using a tangent in the region AB, where the curve becomes approximately straight

and by drawing a line parallel to that through (0, 0) the ultimate skin friction capacities could

be determined as in Figure 3.8.

0

' '

' '

Settlement (mm)

' '

----------1 4----1------_.j! Tangent drawn through

1

' '

I, the selected posit1on 1

L----------~

L,lad (Tons) ' ' ' '

' ' ·' ·'

J'' ~- B I. J~ ---.....,....,,

--------~

i End Bearing 1

L ______ ,

~ - ------ -·l I Skin jj-ictJOil_ J

Figure 3.8 Load vs. Settlement behavior proposed by Van Weele (1957)

Sample piles were selected from the Ceylinco Celestial Residencies (CCR) Project.

So, it should be checked whether the same criteria proposed by Van Weele (195 7) and

Bowles (1996) could be applied to pile tested in Ceylinco Celestial Residencies Project (CCR

project). There are four piles tested under static loading. Only three out of four have passed

the test and those piles were subsequently tested under dynamic loading. For this comparison

those four piles PO 14, P042, P 050 & P078 were used.

P014 is 1200mm diameter pile with a test load of 1260kN, P042 is a 1500mm

diameter pile with a test load of 1875kN and P050 & P078 are 1800mm diameter pile each

vvith test load of 2569.5kN. Figures 3.9, 3.10, 3.11 and 3.12 give the load-settlement

behavior of above mentioned piles under both static loading and dynamic loading. The load

settlement curves obtained from both SL T and PDA test have included in the same graphs for

the easy reference on behavior of those two curves.

Most of the piles tested in Sri Lanka are tested up to 1.5 - 2.0 times working load

rather than going for ultimate loads. In CCR project pile has loaded to 2.5 times working

load.

DESIGN OF CIB PILES

P050 Loads (Tons)

0 250 500 750 1000 1250 1500 1750 2000 2250 2500 2750 3000

2

3

4

5

E 6 E .. 7 r::: G)

E 8 G)

E 9 G)

en 10

11

12

13

14

0 ~

'""' t--

~ '\ -+- Static-Cycle2

l 1'\ - Static-Cycle 1 r-

~ ~

""' ___..,_ PDA t--.....

~~r--, " 1\. ~tangent 1--

"' .._ ~ 1'--. "' r-....

\ ~ ........_ .....,

~ " i"'-.,

"" ~

............. ~~ ~ J

' "' ............... fll... .......... ~ ~ ~

""' ~

b--.~ ~ i' ""' & ""' ['... ~ ~ ~ !'.....

........ ~ !\.. l

.............

'-..... ~ ~

·~---~ ~ 15

Figure 3.9 Application of Van Weele method to the results obtained from P050 Static Load Test

P 042 Loads (Tons)

0 250 500 750 1000 1250 1500 1750 2000 2250 2500 2750 3000 3250

0

1

2

3

4

5

E 6 E 7 .. r::: G) 8 E G)

9 i en 10

11

12

13

14

15

16

~ ~ ~ ,.

~ ~ '-.. \ -+-Static-Cycle2

""" ~ ~ ~ \

1\ ~ -.........::: ~ ""' \ - Static-Cycle

1

"' "~~...... "<::iii

~ [\ ___..,_ PDA ~----___

~ ~ ---""' ~

~ ...... ~ ~"-.... ~ r-.... " "" /_!;

""'i.: t---..... ' ~ ........ I'... l5 1

I I'-. '\ ""' !'-... ~ '--

""" ......._ ... ~ ~ ......

\ I i

: • T L -

'cation of Van Weele method to the results obtained from P 042 Static Load Test

I>ESIGN OF CIB PILES

p 014 Loads (Tons) 0

0

250 500 750 1000 1250 1500 1750 2000

2

3

4

5

E6 E i:7 G)

E .!8 il cng

10

11

12

13

14

~ \ ~ ~ ~ \

\~ ~' \ "" ~ "" \ ..__ "'- ~ \ ~

""' ~ "

I\ -+-Static-Cycle2

\ --Static-Cycle 1

~ \ ___..,._ PDA

' \ ~tangent

~" ~ '~ ~"' ~" "" I'- ' ~ ~ "" ' ~ ~

r--... ' ~

' '\ "" ~~ ~

Figure 3.11 Application of Van Weele method to the results obtained from P 014 Static Load Test

J~'

Skin friction obtained from using PDA results and SLT results are tabulated in Table 3.2.

Table 3.2.Results of the comparison between SLT results with PDA results

Pile No Skin friction from the Skin friction from the Comment

CAPWAP-kN Van Weels method-kN

P 014 (1200 dia.) 9660 3200 Different

P 042 (1500 dia.) 18750 8220 Different

P 050 (1800 dia.) 19450 8200 Different

0

5

10

15

20 E E ~5 G)

E jo G) en 35

40

45

50

0 250 500 750 ~cPO'JI

..... _

---- ·-- --

DESIGN OF CIB PILES

1250 1500 L1~s (Ton~6oo -"\

1\

r\ -+-Static-Cycle2

\ - tangent

_l \

\

\ ----

Figure 3.12 Application of Van Weele method to the results obtained from P 078 Static Load Test

From the Load-Settlement curves given in Figures 3. 9, 3.10 and 3.11, it can see that

there is no 'straight line portion' as described by Van Weele. That means the piles are still in J:;,.

the region where the SF is the governing criteria for the total capacity. This is similar to the

region "OA'' in Van Weele graph in Figure 3.8. For this comparison the line has drawn by

considering the last two point of loading cycle. An assumption was made by assuming those

two points are lying in AB region given in Figure 3.8. Values obtained for skin friction using

Van Weele method are very much on the conservative side and it is not represents the fully

mobilized skin friction. This is because even under the load of 2.5 times working load these

piles were not achieved their full skin friction.

But P078 failed due to "soft toe". It has under gone excessive settlement thus ultimate

skin friction should be mobilized. The application of the Van Weele method to load­

settlement curve of P07 8 yields a mobilized skin friction of 1216 tons. This is very much

high, when compare with other three mobilized SF from SLT given in Table 3.2. So the

expected ultimate skin friction for 1800mm diameter pile is around 1200 tons. This is a

reasonable figure when it compared with the average skin friction of 1800mm pile given in

Table 3.1. Due to the weaker toe area it could not expect SF to be generated in the socketed

DESIGN Of' CIB PILES

area, thus the value of 1216 tons is mainly the SF generated from soil. This ensures the

argument of high SF value developed in the sandy soil region.

Respective PDA test curves are drawn in same figure to make it easy to compare with

SLT Load-Settlement curves. Two curves are moving very closer to each other. This shows

the load and settlement behavior in both static and PDA are very similar in this particular site

used for this case study. This argument again ensures that the results given by the PDA this

case study are acceptable to the required limits of this research.

0

5

10

15

20 E E .. ~25 E Ill E ~30

35

40

45

50

Loading cycle only

0 250 500 750 1000 1250

-!z::,

~ ~

- -

J Loads (Tons)

1500 1750 2000 2250 2500 2750

--r---. r----1 \ -...........

!'+. \ \ -+- P78

1\ -.- P50

\ ..

\ \

-- ~ -·-"---'

Figure 3.13 Load-Settlement behavior of all static load tested piles

The load-settlement behavior obtained from SLT tests are including in the above

figure. Only the loading cycle is considered here. The load-settlement behaviors are very

similar up to 1250MT in the each case except P014 which is a 1200mm diameter pile. P014

curve also acceptable since it shows only a 5mm difference in the settlement up to 1250MT.

As a conclusion it can say that all piles are settled in similar way under different loading and

thus tbe result for SF given by the P078 can be considered as a general case for this case

study.

3.4.3 Discussion on the application of Chin's (1978) method

Chin (1978) proposed that the plot of the ratio between the settlement and the load (SIP) and

the settlement (S) consists of two linear segments. A plot of the (SIP) vs. S for a static load -

settlement curve is shown in the Figure 3.14. According to Chin ( 1978), the inverse of the

slope of the second segment yields the total ultimate carrying capacity and the inverse of the

slope of the first segment gives the ultimate skin friction capacity.

~

0~25 z ~ E ! 0.02

,'

/

~ ~

0~ 15 0 ~

~ =

' ~

ont ] .• E ~ Jl' ' t 0~05 .~-~ ~

0

0 10 20 30 40 Settlement (mm) , ..

Figure 3.14 SIP vs. S considering two linear segments according to the Chin (1978).

Figure 3.15 to 3.17 gives the application of this method for the piles PO 14, P042 and

P050. Tangents to the curves are drawn in a more conservative way by connecting selected points

in the curve. The inverse of the gradient of Line 1 and Line 2 of respective charts are calculated

and tabulated in Table 3.3.

The values predicted by Chin's method match with the CAPWAP results to some extent

and thus show the values given by CAPW AP analysis are acceptable for the purpose of this case

study. Smaller diameter pile tends to give more accurate result than lager diameter piles. This is

obvious as the there are so many other factors involved in calculating the capacities of lager

diameter piles.

I>ESIGN OF CIB PILES

To mobilized ultimate skin friction it should allow the pile to move downwards to some

considerable extent as describe in previous chapters. With out considerable settlement ultimate

capacities would not mobilize. Thus the ultimate total capacities predicted are not match with the

total capacities predicted by the CAPWAP analysis.

Table 3.3 Comparison between SLT results with PDA® results using Chin Method ( 1979)

Pile No

P 014 (1200 dia.)

P 042 (1500 dia.)

P 050 (1800 dia.)

P078 (1800 dia.)

O.OOG

0.0055 --

0.005

' e 1 .!1 0.0045 -~ i e j

().004 .

a.o~s

0.003

a

Skin friction from Skin friction from the

the CAPW AP -MT Chin's method-MT

966 943

1153 923

1945 3458

- -

P 050 - Cnin Mirthod

Lin{} $~~~-·

2 ~ 6 a 1.0

Settlement mm

Figure 3.15 Application of Chin Method for P050 pile

Ultimate

Total

J Capacity-MT

3083

4000

7692

-

iii:

, ..

12 14 1&

I>ESIGN OF CIB PILES

Cl'lln Metod - P 042

0.01)6. - --- --- - --------

o.oos

0.004

t 1 S C.C03< : J ii E

i D.DDJ

O.OOl

D

0 2 4 0 8 10

5il!ltfement mm

Figure 3.16 Application of Chin Method for P042 pile

Chin Metod • P 014 0.011 - · j~·

O.ql :-

0.00!> ..l;.t~d])

'l;: ·~

l ~~ . .... ;.::: -.- L ""' &

~ 0.00!!

I E 0. 1>~7

I ·~· "'

1l.OO!l

o.oos

0.00.:

0 .; 6 " 10 1~

Set:tlr:ment 111m

Figure 3. 17 Application of Chin Method for P014 pile

11

l4

3.5 Estimation of Skin Friction

3.5.1 Beta-Method (~-Method) for Side Resistance

For the analysis of shaft resistance, Johannessen and Bjerrum (1965) and Burland (1973)

established that the unit resistance is proportional to the effective overburden stress in the soil

surrounding the pile. The constant of proportionality is called beta-coefficient, ~' and is

assumed to be a function of the earth pressure coefficient in the soil, Ks, times the soil

internal friction, tan ~1 , and times the quotient of the wall friction (Bozozuk, 1972). Thus the

unit shaft resistance q5 follows the following relations.

q5={3 X av/ {3 = M K5 tanctJ1

Where

Qs - Unit resistance at depth z

~ - Bjerrum- Burland beta coefficient

crv/- Effective over burden pressure at depth z

M- Quotient of wall friction= tan 811 tan ~/

o 1 - Effective soil pile friction angle

~1 - Effective soil friction angle

K5

- Earth pressure coefficient Q

~

rf b

HI qs (/)

Qs , ~

• L-1

Qp

Figure 3.18 Terms and symbols for pile analysis

(01)

(02)

J:;,.

One can develop a wide range of beta-coefficients from a combination of possible earth

pressure coefficients, friction angles and wall friction quotients. However it appears that the

variation of the beta-coefficient is smaller than the variation of its parts would suggest.

In analyzing measurements on piles subjected to down drag, Bjerrum et al. (1969) found that

the beta-coefficient in the soft silt clay ranged between 0.20-0.30. This range can be

considered the lower boundary of the beta-coefficient. While the theoretical upper boundary

obviously can be very large, there is a practical limit governed by the density and strength of

the soil in which the pile is driven or otherwise installed. For piles in very dense soil the

upper boundary can be approach and exceed a value of 1.0, but usUally an upper limit of 0.8

is assumed. Following table suggests a relative range of beta values. The ranges shown are

very wide and very approximate.

Table 3. 4 range of Beta -Coefficients

Soil Type ~ beta

Clay 25-30 0.23-0.40

Silt 28-34 0.27-0.50

Sand 32-40 0.30-0.80

Gravel 35-45 0.35-0.80

,~.

Although it has been proven conclusively that the transfer of load from a pile to the soil by

means of shaft resistance is governed by the effective stress, for piles in clay, a total stress

analysis can be useful in site-specific instances. Also, enough information is often not

available to support a reliable design based on effective stress analysis. A total stress analysis

may then be used, which means that the shaft resistance is equal to the untrained shear

strength of the soil and independent ofthe overburden stress:

qs = aru

Where

Tu- Untrained shear strength

a- Proportionality coefficient

(03)

However, the total stress analysis can only lead so far and effective stress analysis according

to equation 01 provide better means for analysis of test data and for putting experience to use

in a design. Of course, more sophisticated effective stress theories for unit shaft resistance

exist. However, in contrast to most of these, the effective stress approach according to

Equations (OJ) is not restricted homogeneous soils, but applies equally well to piles in

layered soils and it can easily accommodate non-hydrostatic pore pressures.

Equation (04) gives the total shaft resistance as the integral of the unit shaft resistance over

the embedment depth:

Q5 = f0h rz dz = f

0h As (cl + {3crvi/ )dz (04)

Where

Qs- Total shaft resistance (fully mobilized)

As - pile unit circumferential area

h- Pile embedment depth

For sandy soil where c1 tends to zero, the above equation can be modified as follows.

Q5 = f0h rz dz = f

0h As {3crvil dz (05)

Das (1999) has followed the same principals and derived the much simplified versions of

above equations for side resistance in sand. The method by Das (1999) is based on at-rest

earth pressure coefficient,K0 the average effective vertical stress found at the midpoint of the

soillayer,av/ and the friction angle ~· The total side resistance in cohesion less soils is found

by Equation (05), where the effective stress, av/ is multiplied by its pertaining empirical J:;, .

beta factor, ~' given in Equation (06) & (07), and the depth of the soil layer, D. The

summation of this product from each layer multiplied by the perimeter length gives the total

side resistance,Qs.

Qs = P X D X S X ({3 X av/) X Li (06)

f3 = K0 x tan~ (07)

K0 = 1- sin~ (08)

3.6.1. ICTAD Method (ICTAD Specification for Pile Construction)

This is one of the methods, which is most commonly used in Sri Lanka. According to

this method skin friction totally depends on the SPT N values. Variation of skin friction

along the pile shaft is similar to the variation of SPT N values. This is one of the simplest

methods that can be used to evaluate skin friction of bored piles.

Skin friction up to the Rock,

Q5 = 1.3 X Ab X N

Skin friction in the rock

Q5 = 2.0 X Ab X N

Sample calculation is given under Annexure A.3.2.1.

3.6.2. O'Neill and Reese (1999)

(09)

(9a)

Methods used to evaluate skin friction in this section are extracted from the research project

conducted by David A.Seavey and Scott A. Ashford in the Department of Structural

Engineering of University of California in December 2004. Discussion in this section is

mainly to give a basic idea on these methods in a structural point of view and readers are

referred to full research report "Effect of Construction Methods on the Axial Capacity of

Drilled shafts" by same authors for further details. ,~.

The beta-method given in O'Neill and Reese (1999) is one of the methods that most

commonly used in practice. It gives the following equations for finding the unit side

resistance, q5 (kPa) and ~' where ~ is the beta factor for the pertaining layer.

qs = f3 x <Iv/ (10)

The beta factor is a dimensionless correlation factor between the vertical effective

stresses, crv/, found at the midpoint of the soil layer, and the unit side resistance, q5• Beta is

limited to a minimum of 0.25 and a maximum value of 1.20 (0.25 ~ ~i ~1.20) and q5 must

not exceed 200 kPa (2.1 tsf) (O'Neill and Reese 1999 and Caltrans 2000).

For sands with an SPT N-value greater than or equal to 15 (N 2: 15), ~i is found by

(11) in metric units, where SPT N is the average SPT blow count for the soil layer, and Zi 1s

the vertical distance from the ground surface to the middle of the soil layer, in meters.

= 1.5- 0.245 x Cza 0·5 11)

If the SPT N-value is less than 15 (SPT N < 15) then the dimensionless correlation

factor is scaled by a ratio of the SPT N-value.

f3i = (:S) (1.5- 0.245 X (zJO.S) (12)

For gravelly sands or gravels with an SPT N-value greater than I5, O'Neill and Reese

(1999) provide the following Equation (13). However, if the SPT N-value is less than 15, ~i

is scaled accordingly as shown with Equation (14).

{3i = 2.0- 0.15 x (zJ0·75 (13)

f3i = (:S) (2.0- 0.245 X (zi) 0·5) (14)

Soil that exceeds a blow count of 50 is named as an intermediate geomaterial (IGM). The

following equations apply for the side resistance of IGMs in cohesion less soils.

q 5 = av/ X Koi X tan~~ (15)

0.34 ):;,.

~~ = tan- 1 N1 (16)

( 123+203x( :0))

( ) sin<)>~ Koi = (1- sin~/) X 0.2 X PaX N/

(JVl

(17)

The blow count value, SPT N, should be limited to 100, even if tests give a higher

value. The angle of internal friction,~~ pertaining to the layer of consideration, and Kai is the

at rest earth pressure coefficient in the layer. The vertical effective stress ati is found at the

midpoint of the layer. Sample calculation is given Annexure A.3.2.2. In this sample

calculation one pile has selected (P088) and layer by layer evaluation has done using O'Neill

and Reese method (1999).

3.6.3. Vesic Method (1967)

Most of the design methods developed for piling are empirical methods. But the

Vesic method developed in 1967 has developed based on the behavior of the vertical stress

around the pile. In order to develop a method of ultimate load prediction that better

represents the physical reality than the other approaches, and is not excessively complicated,

an idealized distribution of effective vertical stress O"ti with depth adjacent to a pile is shown

in Figure 3.19. O"ti is assume to be equal to the overburden pressure to some critical value

(CJti) at a critical depth Zc, beyond which vertical effective stress remains consistent. The use

of this idealized distribution, although simplified, leads to the two main characteristics of

behavior observed by Vesic: namely, that the average ultimate skin resistance and the

ultimate base resistance become constant beyond a certain depth of penetration. If the pile­

soil adhesion Ca and the term cNc are taken as zero in equation and the term

~ {JByNy neglected as being small in relation to the term involving Nq, the ultimate load

capacity of a single pile in sand may be expressed as follows.

Pu = J Fw CO"tiKs tan~a dz + AbO"tbNq- W (18)

m v m :.~ J Z

J~'

L o'vc

j Figure 3.19 Simplified distributions of vertical stresses adjacent to pile in sand

On the basis of the test results of Vesic (1967), values of the dimension less critical

depth Zc/d and K5 tan~~ with ~/,are shown in Figures 3.21 and 3.20 respectively.

-o

' u ...

z c /d vs (6

2Qa I J I I I I t t t I t i I i l

i5 1--l

~~ ~£\ A.~

0~

Figure 3.20 Charts of Zcvs. ¢/

In a layered-soil profile, the critical depth Zc refers to the position of the pile

embedded in the sand. It should be emphasized that these relationships may be subject to

amendment in the light of further test results. For example, at present, the dependence of the

K5 tan~~ on the pile material is not defined. Vesic' s tests were carried out on steel tube piles,

but the values of K5 tan~/ derived from these tests appear to be applicable to other pile

a

materials. However, in the light of future test results it may be possible to derive different

relationships for different pile materials.

The values for K5 tan~/ can be obtained from the data of Meyerhof (1976) and the

a

graph is shown in Figure 3.21. The values given in the graphs are in the range of32°::::; ~/::::;

38°. But some of the ~/ values obtained from the soil tests conducted in local sites are much

higher than that as shown in Table 3.5. So, there is a restriction to use above chart to obtain

K5 tan~~ values for soil with high ¢/value.

Table 3.5 practical valuesfor ¢/obtained from local projects.

Project ~I

Dawson Grand Tower, at Colombo 02. 28°-50\J

Trillium Residencies, Colombo 08 30u -50°

Mayfair City, Colombo 03 30°-50\J

I

Volwzs of K 5 ton 0 0 Bosad on Ma:yc:zrhOf ( 1976)

1·6 tt J t t I 1 tIft 1 lIt 11

t. 2

0·8

-- 0 ~ c 0

0-41 I ~ .. ,_-...........,..........-{

I l I I I. I li Q I I I I l I I ! ~0 40

0' 1

Figure 3.21 Mayerhof(l976) charts ofK5 tan rfa vs. ¢/

3.8.4. William et al. (1981)

William et al. ( 1981) suggested that for piles installed in sandstone, mud stone or

shale. the skin friction qs in the rock socketed length of the pile can be estimated using the

relationship given in Equation ( 19). This method can only be used to estimate skin friction

generate in weathered or intact rock layers.

fs = af3quc (19)

que is the unconfined compression strength of socketed rock, which can be calculated using

point load index given in rock compressive test results.

Where the factors a and ~can be obtained from the Figure 3.22 & Figure 3.23. The

mass factor 'j' can be obtained from the guidelines proposed by Hobbs ( 1975) given in Table

3.6

Table 3. 6 Estimation olmassfactor .'i'

RQD (%) 0-25 25-50 50-75 75-90 90-100

Fracture Frequency per meter Mass factor ''j"

15 0.2

15-8 0.2

8-5 0.2- 0.5

5-1 0.5- 0.8

I 0.8 -1.0

1.0 11 i > I i t I I I I I I I t I f J I I f -r 1 1 1

rn

0.8

0

~ 0.6

.;:! c; 0

"fi ~ "' ~ 0.4

"' " 0

"' "' " &. 0.2

0.0

'" e<fJqK ~ \ I

)J.....----Rosenberg & Jotrmeaux

Unconfined compression strength, que MN/m2

Reduction factors for rock socket skin friction ( Tomlinson, 1994)

Figure 3.22 Chart for rock socket correctionfactor, a

I ! ' I

..

1.0

00 0.8

B-u co -r: 0.6 0 u (!) ,_ 0 u 0.4 0)

""' (,)

0 If)

t5 0,2 0

0:::

0

Pile bearing on rock

0.2 0.4 0.6 0.8

Mass tactor,J

Figure 3.23 Chartfor rock socket correctionfactor, fJ

1.0

The relationship given by equation assumes that the side of the rock socket is free

from debris and not smeared with the material wash down the borehole. A conservative

analysis, assuming RQD less than 25% and unconfined strength of 10 MPa, yield a and ~

values 0.65 and 0.13 respectively. The use of the relationship given in the equation results the

skin friction resistance of 845 kPa within the rock socket. Wyllie (1991) suggested that the

use of bentonite during the drilling process reduces the skin friction and if bentonite is used

the skin friction of the rock socket should be taken as 25% of a clean rock socket.

3.9. Tip Resistance (End Bearing)

When designing for cohesionless soil conditions under the axial loading of bored and

cast in-situ concrete piles, soil investigations must determine what the density of the soil is,

usually by a standard penetration test (SPT). The load-bearing capacity is then calculated

based on the SPT blow count. N. Blow counts are the number of times it takes for a dropped

hammer to penetrate one foot into the soil, which directly relates to the bearing resistance.

However, since the bored and cast in-situ concrete piles are stationary objects that bear on

the soil in a less dynamic manner than a dropped hammer, the blow count value can be

increased by a specified amount to obtain the estimated bearing capacity. According to

Caltrans, soil is considered to be competent when theN-value is greater than or equal to 20

for upper layers, and 30 for lower layers. This means that the soil is adequate enough to

I I,

withstand axial loads without remediation of the soil. Soil is considered poor when the N­

value is less than 10, and the soil is too weak to withstand axial loading. When the value is

between 10 and 20, the soil is classified as marginal and additional investigation is

recommended (Caltrans 1997).

According to O'Neill and Reese (1999), the soil is classified as cohesionless when

the blow Count is less than or equal to 50. and the tip resistance can be found by Equations

(20) & (21). If SPT value is greater than 50, the material is an IGM.

qp = 0.60 X N (tsf) (20)

qp = 57.5 X N (kPa) (21)

This procedure is fairly simple in application and thus should be used as an estimate

of the tip resistance and not taken an exact value. Caltrans uses this equation for

cohesionless soils with N-values less than 75, and qP is limited to a maximum value of 3830

kPa (40 tst}

For IGMs the following Equation (22) is recommended by O'Neill and Rees.e (1999).

Other equations for reduced base resistance for IGMs can be found in the 1999 FHW A

Manual for Drilled Shafts.

qp ( )

0.8

0.59 x ~ x a~b(kpa) Palavb

(22)

As with the side resistance IGM equation, N should be limited to 100. The

atmospheric pressure,Pa in the Sl system is 1 01 kPa, and the vertical effective stress, a~b is

the value calculated at the elevation at the base of the pile, in kilopascals (O'Neill and Reese

1999). Similar to calculating the bearing capacity in clays, displacement limits may also be

taken into account when calculating the bearing resistance in sands. Das (1998) recommends

reducing the above equation to qpr when the pile diameter, D, exceeds 50 inches. Caltrans

also uses this reduction

- [~] xq Qpr - D(in) p (23)

The lip resistance may also be calculated using the effective stress at the base of the pile

multiplied by a bearing capacity factor,Nq, which is difficult to obtain in the case of CIB

piles because the original Nq factors were based on piles, which are driven, and CIB piles are

not. Therefore, Nq values are lower than what is used in calculations for bearing capacities of

piles Several researchers have provided values for Nq for CIB piles; however, these values

tend to vary by a great amount, so experience and good engineering judgment must be

utilized. The weight of the CIB pile is subtracted out of this equation by assuming that the

em pile weight is approximately equal to the soil it has replaced.

Qp,net = Ab X (J~b X (Nq - 1) (24)

Values for N q can be found in tables provided in textbooks under methods for Vesic,

Meyerho[ and Terzaghi, to name a few (e.g. Das 1999).

ln general, tip capacity is mobilized at displacements that far exceed the displacements

required to activate side resistance (Osterberg 2000). This can be attributed to the

construction of the pile. Disturbed or loose soil at the tip of the pile, due to excavation or

drilling t1uid deposits, must compact before it will provide resistance. This requires several

inches of displacement. If this sediment could be compacted beforehand, this required

displacement is eliminated and axial capacity can be mobilized to resist loads in unison with

the side resistance (e.g. Osterberg 2000, Walter eta!. 2000, Littlechild et al. 2000, Dapp et

a!. 2002, and Mullins et al. 2000). For example, Caltrans will only recognize full tip

capacity to be contributing after a displacement equal to 5% of the pile diameter has

occurred, and typically, tip resistance is completely discounted due to the possibility of a

soft bottom occurring. Cleanout methods, such as pressure washing with U-tubes (Lin,

2000), and post-grouting methods can aid in preventing soft bottoms. Studies concerning

post-grouting have proven to be very effective in reducing displacements and increasing the

axial capacity. The methods covered so far are for the axial design of CIB piles, which does not take

into account the reinforcement design. CIB piles are also designed structurally in order to

\Vithstand t1exural (lateral) and axial loads. Lateral behavior is not the focus of this research

study.