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DETERMINING SLIPPING STRESSOF PRESTRESSING STRANDSIN PILES EMBEDDEDIN CIP CAPS

by

Mohamed Khaled ElBatanouny

Bachelor of ScienceHelwan University, 2008

______________________________________________

Submitted in Partial Fulfillment of the Requirements

For the Degree of Masters of Science in

Civil Engineering

University of South Carolina

2010

Accepted by:

Dr. Paul Ziehl, Advisor

Dr. Juan Caicedo, Committee Member

Dr. Fabio Matta, Committee Member

Tim Mousseau, Dean of The Graduate School

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Copyright by Mohamed K. ElBatanouny, 2010All Rights Reserved.

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DEDICATION

To my family, my fiance and friends whose love and support helped me along

the way. To my professors at U.SC and Helwan University, Egypt.

I would like to give special thanks to my parents, General Khaled ElBatanouny

(May God shower his soul with mercy) and Prof. Fatma Shaltout, my sisters (Hadeer and

Alaa), my brothers (AlHussien and Abdel-Rahman) and my fiance (Marwa) and her

family (specially her father Prof. Ahmed Mousa). Although they are thousands of miles

away, their emotional support during my study always encourages me and pushes me to

achieve my goal and finish my study.

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ACKNOWLEDGEMENTS

I would like to thank my advisor Dr. Ziehl for his time, support and guidance

during the research. I would like to thank Dr. Caicedo for his help in the lab and with

data acquisition and Dr. Matta for his support as a member of my defense panel. I would

like to thank the SCDOT and FHWAfor financing this research work.

I would like to thank Dr. Timothy Mays for his input in the theoretical models and Avery

Fox for supervising of the structures lab. Also, I would like to thank my fellow graduate

students Aaron Larosche, Shawn Sweigart and Jese Mangual for their assistance in lab

testing as well as all the undergraduates that put in hard work alongside us. The input of

Dr. Jose Restrepo related to the experimental program is also gratefully acknowledged.

Finally, I would like to thank the support staff in the Department of Civil and

Environmental Engineering that helped guide me through this process.

First and last, great thanks are due to God ALLAH, without his will nothing can

be achieved.

ABSTRACT

A lateral cyclic loading test has been completed involving four prestressed concrete pile

specimens embedded into cast in-place (CIP) reinforced concrete bent-caps. The study

was conducted to determine the moment capacity of the piles which is significantly

affected by the slipping stress of prestressing strands. In general the ACI 318 Eqn. (12-4)

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is used to calculate the development length and slipping stress of prestressing strands, yet

the validity of this equation for strands in confined sections is not proven. The piles used

were 18 in. (457 mm) square x 18 ft (5.5 m) long and were plainly embedded into the

bent-cap to depths of 18 in. (457 mm) and 26 in. (660 mm). Analytical models were

created in which the slipping stress of the strands was calculated using ACI code equation

for development length. In all cases the connection achieved more moment than expected

due to the confining stress produced from the bent-cap which tends to decrease the

development length of the strands. A modified equation that accounts for the beneficial

effect of concrete confinement due to shrinkage is introduced.

Experimental data from a research completed in 1992 was used to validate the

proposed equation. Analytical models were created in which the slipping stress values

were calculated using the proposed equation, the current ACI code equation, and the

experimental values. Generally, the proposed equation has a better match with the

experimental results than the ACI code equation.

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TABLEOF CONTENTS

DEDICATION...........................................................................................................................iii

ACKNOWLEDGEMENTS..............................................................................................................iv

ABSTRACT...............................................................................................................................v

LISTOF TABLES......................................................................................................................ix

LISTOF FIGURES......................................................................................................................x

INTRODUCTION.........................................................................................................................1

1.1 BACKGROUND..........................................................................................................1

1.2 R ESEARCH SIGNIFICANCE...........................................................................................3

1.3 OBJECTIVES.............................................................................................................4

1.4 LAYOUTOF THESIS...................................................................................................4

LITERATURE REVIEW................................................................................................................6

2.1 INTRODUCTION.........................................................................................................6

2.2 HISTORYOFTHE DEVELOPMENT LENGTH EQUATIONFORPRESTRESSING STRANDS.............6

2.3 PILESTO CAST-IN-PLACE CAPS CONNECTIONS...........................................................14

BENT-CAP CONFINING STRESS EFFECTONTHE SLIPOF PRESTRESSING STRANDS............................20

3.1 ABSTRACT.............................................................................................................21

3.2 INTRODUCTION.......................................................................................................21

3.3 R ESEARCH SIGNIFICANCE.........................................................................................23

3.4 EXPERIMENTAL PROCEDURE.....................................................................................23

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3.5 MOMENT CURVATURE MODELS...............................................................................27

3.6 DEVELOPMENT LENGTHAND CONFINING STRESS........................................................28

3.7 PROPOSED MODIFICATIONSTO EXISTING ACI EQUATION..............................................33

3.8 R ESULTSAND DISCUSSION.......................................................................................34

3.9 SUMMARYAND CONCLUSIONS..................................................................................37

3.10 ACKNOWLEDGEMENTS.............................................................................................39

3.11 R EFERENCES..........................................................................................................39

DETERMINING SLIPPING STRESSOF PRESTRESSING STRANDSIN CONFINED SECTIONS.......................50

4.1 ABSTRACT.............................................................................................................51

4.2 INTRODUCTION.......................................................................................................51

4.3 R ESEARCH SIGNIFICANCE.........................................................................................53

4.4 SUMMARYOF SHAHAWYAND ISSA3 TEST RESULTS.....................................................53

4.5 HISTORYOF EXISTING ACI EQUATION.......................................................................57

4.6 MOMENT-CURVATURE ANALYSIS.............................................................................60

4.7 R ESULTSAND DISCUSSION.......................................................................................62

4.8 SUMMARYAND CONCLUSIONS..................................................................................63

4.9 ACKNOWLEDGEMENTS.............................................................................................64

4.10 R EFERENCES..........................................................................................................65

SUMMARYANDCONCLUSIONS...................................................................................................74

5.1 SUMMARY.............................................................................................................74

5.2 CONCLUSIONS........................................................................................................75

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5.3 R ECOMMENDATIONSAND FUTURE WORK...................................................................76

REFERENCES..........................................................................................................................77

APPENDIX A.........................................................................................................................80

APPENDIX B..........................................................................................................................87

APPENDIX B

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LISTOF TABLES

Table 3.1: Maximum moment and slipping stress (measured and current ACI Equation).

...........................................................................................................................................42

Table 3.2: Comparison of calculated slipping stress (current ACI Eqn. and proposed

Eqn.)...................................................................................................................................42

Table 3.3: Comparison of slipping stress and maximum moment....................................42

Table 4.1: Details of test program [Shahawy and Issa].....................................................67

Table 4.2: Shahawy and Issa test results...........................................................................67

Table 4.3: Comparison of theoretical development length................................................68

Table 4.4: Relation between embedment length and confinement....................................68

Table 4.5: Results of slipping stresses...............................................................................69

Table 4.6: Ultimate moments for different slipping stresses using moment-curvature

analysis..............................................................................................................................70

Table 4.7: Required Development length to achieve experimental slipping stress...........70

TABLE 4.7: REQUIRED DEVELOPMENTLENGTHTOACHIEVEEXPERIMENTAL

SLIPPINGSTRESS

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LISTOF FIGURES

Figure 3.1: Prestressed pile reinforcing details and strand arrangement...........................43

Figure 3.2(a): Overview for the test setup with hydraulic actuator...................................43

Figure 3.3: Closeup of the LVDTs (encased in steel tubes) for measuring strand slip....44

Figure 3.4: Schematic of the test setup..............................................................................45

Figure 3.5: Pile/bent-cap confining mechanism................................................................45

Figure 3.6(a): Crack and spalling on top face of pile with strand visible..........................46

Figure 3.7: Strand end slippage vs. moment for BC-26-1.................................................47

Figure 3.8: Effect of confinement on development length................................................47

Figure 3.9: Moment vs. displacement for BC-18-1...........................................................48

Figure 3.10: Moment vs. displacement for BC-18-2.........................................................48

Figure 3.11: Moment vs. displacement for BC-26-1.........................................................49

Figure 3.10: Moment vs. displacement for BC-18-2.Figure 3.11: Moment vs.

displacement for BC-26-1.

Figure 4.1: Details of test specimens.................................................................................71

Figure 4.2: Moment vs. curvature for A-2E......................................................................71

Figure 4.3: Moment vs. curvature for B-4E......................................................................72

Figure 4.4: Moment vs. curvature for C-5I........................................................................72

Figure 4.5: Moment vs. curvature for D-2I.......................................................................73

Figure 4.6: Average ultimate moment vs. embedment depth............................................73

FIGURE 4.6: AVERAGEULTIMATEMOMENTVS. EMBEDMENTDEPTH.

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CHAPTER I

INTRODUCTION

1.1 BACKGROUND

According to the United States Geological Survey report [USGS, 2009], South Carolina

is the most seismically active state in the eastern United States. A major earthquake hit

Charleston in 1886. This earthquake was estimated to be a 7.3 on the Richter scale. Ten

earthquakes occurred between the years 1974-2003 and registered 3.5 or higher on the

Richter scale. Recently an earthquake hit Summerville, SC and registered 3.6 on the

Richter scale. The New Madrid seismic zone which is an active seismic zone centered in

southeastern Missouri also affects the northwest corner of the state. Based on the above,

it is required for all designed structures to endure design earthquakes without failure. Of

particular interest are highway bridges which play a vital part in the transportation of road

and rail vehicles. Minimizing damage and preventing the collapse is important to these

structures as they facilitate the evacuation of damaged urban areas and the movement of

emergency vehicles [Sweigart, 2010].

Precast prestressed piles are commonly used in the bridge construction industry in

South Carolina and throughout the United States, often in combination with cast in-place

(CIP) reinforced concrete bent-caps. However, the performance of the connections

between prestressed piles and the CIP bent-caps under seismic loading is not well

understood. The SCDOT Seismic Design Specification for Highway Bridges [SCDOT,

2008] is used by the SCDOT in conjunction with the AASHTO LRFD Bridge

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Design Specifications [AASHTO, 2004] to design bridge structures to meet the

demands of seismic events.

The current SCDOT standard connection detail utilizes a plainly embedded

prestressed pile having embedment distance of one pile diameter with a construction

tolerance of 6 inches (152 mm). Plain embedment requires no special detailing to the

pile end or the embedment region and no special treatment of the pile surface such as

roughening or grooving. This connection detail is very cost effective as it requires no

special equipment or time consuming labor processes in the field. However, the ductility

and moment capacity of such connections is not well understood. The short embedment

length is often much less than the length required for developing the full capacity of the

prestressing strand. Given that South Carolina is the most seismically active state in the

eastern United States, this investigation has been conducted to determine the structural

performance of such connections under reverse cyclic loading.

The slipping stress and development length of prestressing strands are generally

calculated from ACI 318 Eqn. (12-4)[ACI 318, 2008], yet the validity of this equation for

strands in confined sections is not proven. This equation was introduced for the case of

superstructure elements which are not necessarily subjected to confining stress. On the

other hand, a pile embedded into a bent-cap is subjected to the shrinkage of the confining

concrete in the bent-cap which creates confining stress (also known as clamping force)

on the pile which serves to enhance the bond between the prestressing strand and the

surrounding concrete leading to a decrease in the development and increase in the

slipping stress of the prestressing strand.

In an article titled, Effect of Pile Embedment on the Development Length of

Prestressing Strands [Shahawy and Issa, 1992], the authors discuss the findings of their

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research related to precast prestressed concrete piles embedded in reinforced concrete

bent-caps under lateral loading. Due to the shrinkage of the bent-caps a confining stress

affects the pile and tends to decrease the development length of the prestressing strands.

Using the ACI code equation in such conditions leads to a conservative value of required

embedment depth of the pile in the bent-cap to achieve the maximum capacity of the

prestressing strands.

1.2 R ESEARCH SIGNIFICANCE

In the program described herein, full-scale precast prestressed concrete piles were

embedded in CIP bent-caps and were tested at the University of South Carolina

Structures Laboratory. The results were compared to a theoretical moment-curvature

analysis by setting the slipping stress value using the current ACI code equation. The

experimental and theoretical moments did not agree. An equation for calculating the

confining stress exerted from shrinkage of the bent-caps on the piles is therefore

introduced. This equation was used to modify the current ACI code equation by adding

the beneficial effect of confinement. Confining stress tends to enhance the bond between

concrete and prestressing strands and this enhancement in bond leads to a significant

decrease in the required development length for the strand to reach its capacity. In the

case of piles embedded in CIP caps for a known distance the enhancement in bond leads

to an increase in the slipping stress of prestressing strands. The modified ACI equation

was then used for modeling the piles. These models resulted in a better match when

compared to experimental results.

The modified ACI code equation and the current ACI code equation were then used

to run analytical models for the experimental work done by Shahawy and Issa (1992) and

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the results compared. The modified code equation approach resulted in a better match

with the experimental results.

1.3 OBJECTIVES

The main objective of this study was to determine a proper equation or analytical model

to calculate the slipping stress of prestressing strands for piles embedded in CIP caps.

This is critical because prestressing strands in CIP caps are affected by a confining stress

due to shrinkage in the concrete surrounding the pile. Specific objectives of this study can

be summarized as follows:

1. Determine the validity of ACI code equation for development length of

prestressing strands for CIP caps.

2. Calculate the confining stress exerted from the shrinkage of CIP caps on piles.

3. Develop an equation to calculate the development length of prestressing strands in

confined sections.

4. Compare the results from the current ACI code equation and the proposed

equation with experimental results.

1.1 LAYOUTOF THESIS

The Thesis consists of five chapters. In Chapter II, background information about

development length and the history of the current equation in ACI 318 are discussed. A

summary for some of the studies which are relevant to this case study is presented.

Chapter III and Chapter IV were written in paper form and submitted for

publication as journal articles. Therefore, description of some findings and basic

information regarding development length are repeated in some cases.

Chapter III is titled Bent-cap Confining Stress Effect on the Slip of Prestressing

Strands, where the experimental study conducted in the U.SC structures lab is presented

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and compared to the current ACI code equation. An equation for calculating confining

stress is proposed in this chapter and used to modify the current ACI code equation for

development length of prestressing strands.

Chapter IV is titled Determining Slipping Stress of Prestressing Strands in

Confined Sections. In this chapter the modified ACI code equation and the current ACI

code equation were used to model pile specimens tested in 1992 under the effect of

confining stress. The experimental results provided from the study conducted in 1992

were used to validate the modified equation compared to the current ACI code equation.

Chapter V includes a summary of the Thesis as well as conclusions based on this

study. Recommendations for further research are also provided in this chapter.

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CHAPTERII

LITERATURE REVIEW

1

2

2.1 INTRODUCTION

Numerous studies were conducted on the development length of prestressing strands in

the 1950s and the 1960s. These investigations were performed to understand the nature

of the bond between concrete and prestressing strands. These studies lead to the current

ACI code approach which addresses the development length by dividing it into two

distances with different bond characterization; transfer length and flexural bond length.

Unfortunately, the current ACI equation does not address conditions where confining

stress is present. However, recent investigations have been conducted to examine pile to

cast in place (CIP) bent-cap connections and subsequent development length of

prestressing strands. A brief review of the history of development length and the reasons

for developing a new equation to address the effect of confining stress is presented in this

chapter.

2.2 HISTORYOFTHE DEVELOPMENT LENGTH EQUATIONFORPRESTRESSING STRANDS

Janney (1954), Nature of Bond in Pre-Tensioned Prestressed Concrete

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Janney [1954] conducted this study to identify the types of bond present within a

prestressed concrete section to prevent the slipping of the prestressing strands. Janney

used two types of tests to define transfer bond stress, flexural bond stress, and the

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interrelation between them. The presence of two bond stresses was referred to the fact

that the prestressing tension in the strand is transferred to concrete by bond (transfer

bond) while after deformation due to flexure a bond is formed with the same mechanism

of the bond which forms in ordinary un-prestressed sections (flexural bond).

In order to determine transfer bond, Janney used prestressed concrete prisms

reinforced with either wire strands or one strand while using different concrete strength

and surface roughness. The second test was performed to determine flexural bond stress

by testing short beam specimens to failure while monitoring the steel strain throughout

the length of the beam. Wire strands or one strand was used for prestressing to different

pre-tension levels. The effect of surface roughness was also examined.

Janney reported that there are three main factors that contribute to the transfer

bond stress: adhesion between concrete and steel, friction between concrete and steel, and

mechanical resistance. The largest contributor for transferring stress to concrete is friction

as mechanical resistance can be neglected due to lack of deformations in prestressing

strands while adhesion is destroyed due to relative slipping at the release of pre-

tensioning stress. An elastic, thick walled cylinder analysis was used to calculate the

concrete stresses surrounding the tendon within the transfer length. The maximum radial

compressive stresses and circumferential tensile stresses were 3,300 psi (23 MPa) at a

prestressing level of 120,000 psi (825 MPa). This value exceeds the elastic properties

assumed; therefore true elastic action is not expected. In spite of the high computed

tensile stresses, the investigation of the concrete at the ends of the prisms did not show a

failure in the cement paste, therefore it is believed that sufficient inelastic yielding took

place to relieve these tensile stresses. Conclusions derived from the first test were as

follows:

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Effect of wire diameter on the transfer bond of strands was mentioned;

when diameter increases the transfer length tends to increase.

Change in concrete strength is not expected to affect the coefficient of

friction between steel and concrete, therefore the transfer bond is

independent of concrete strength.

Surface condition has a more significant effect on transfer bond. For

example in the case of rusted wires the full pre-tension stress is

transferred to concrete more rapidly and in less distance than

corresponding clean wires.

The second test was designed to examine flexural bond stress. The bond failure

occurs when flexural bond stress concentrations reach the end region of a prestressed

beam (i.e. when the flexural bond stress overlaps the transfer bond stress). As the flexural

bond stress appears only after flexural cracking, this wave progresses smoothly ahead of

the flexural crack from the center to the end of the beam. The experimental results show

that beams with clean wires failed in bond while beams with rusted wires failed by

fracture of the wires. The highest average flexural bond stresses at bond failure for beams

with clean wires varied from 160 to 220 psi (1.1 to 1.5 MPa) while for beams with rusted

wire this range was from 500 to 800 psi (3.4 to 5.5 MPa). Yet corroding strands might

result in a decrease in the cross section of strands or localized corrosion leading to

unexpected failure of the section. The validity of the expression used for calculating

flexural bond stress in ordinary reinforced concrete was examined. It was concluded that

this expression does not yield accurate results for the case of prestressed concrete. The

main conclusion was that the flexural bond length is approximately equal to the

difference between the total length and the transfer length.

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Hanson and Kaar (1959), Flexural Bond Tests of Pretensioned Prestressed Beams

Hanson and Kaar [1959] conducted this investigation to determine flexural bond stresses

in beams pre-tensioned with seven-wire strands. Their main variables were the

embedment depth and the diameter of strand. Forty-seven simply supported beams were

tested throughout the study. The beams were divided in groups having different numbers

of specimens in order to evaluate; a) the effect of different strand diameter and

embedment length for beams having different shearspan ratio; b) the effect of concrete

strength on bond performance; c) the effect of reinforcement percentage on the bond

stress; and d) the effect of surface conditions on bond performance. They used three

different strand diameters of , , or in. (6.4, 9.5, or 12.7 mm) throughout the study.

The effective prestressing of the strands used had a range starting from 120 to 147 ksi

(828 to 1014 MPa). A variety of steel strands were used with an ultimate strength

exceeding 250 ksi (1724 MPa). The moment at cracking, the moment at initial end slip,

and the moment at ultimate strength were evaluated during the test while the flexural

ultimate moment was calculated. It was realized that with a longer embedment length the

ratio between measured moment and calculated flexural ultimate moment increases.

The average bond stress over the entire embedment depth was calculated. It was

stated that this average bond stress is a convenient measurement of the bond

performance of prestressing strands. At longer embedment lengths, the resulted flexural

wave form was found to include a longer area from a peak near the transfer length to the

section of maximum steel stress. Therefore, the average bond stress just before general

bond slipping decreases with longer embedment lengths. For the strand to withstand

general bond slipping, embedment lengths were suggested to be 70, 106, and 134 in. (1.8,

2.7, and 3.4 m) for strand diameters , , or in. (6.4, 9.5, or 12.7 mm), respectively. It

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was also mentioned that rusted wires had better bonding properties than clean wires

bearing in mind that localized pitting might occur if strands are corroded.

In general, Hanson and Kaar agreed with Janneys flexural bond wave theory

confirming that general bond slipping occurs when the peak of the high bond stress wave

reaches the prestress transfer zone. As when the flexural bond wave reaches the prestress

transfer zone, the steel stress increases followed by a decrease in the strand diameter.

Consequently, the frictional resistance is reduced and general bond slipping occurs.

Kaar, LaFraugh and Mass (1963), Influence of Concrete Strength on Strand

Transfer Length

Kaar, LaFraugh and Mass [1963] conducted this study to investigate the effect of various

concrete strengths on the transfer length of prestressing strands. The test consisted of

thirty-six concrete prisms, resulting in a total of 72 transfer lengths with the strand

diameter and effective prestressing stress as variables. The companion transfer lengths

were averaged to give a total of 36 transfer lengths. One mix design was used through all

specimens where the strength at transfer was varied by varying the time of releasing of

the strands from one to thirty days. The compressive strengths of concrete varied from

1,660 to 5,000 psi (11 to 34 MPa). Different strand diameters were employed during the

tests of , , , and 0.6 in. (6.4, 9.5, 12.7 and 15.2 mm). The prestressing force of the

strands ranged from 146 to 180 ksi (1010 to 1240 MPa). Transfer lengths were evaluated

by measuring concrete surface strain with a Whittemore mechanical strain gauge.

Measurements were performed at ten time intervals starting immediately after transfer till

an age of one year to evaluate the effect of time on transfer length.

The variation of concrete strength did not have significant effect on the transfer

lengths of strands up to in. (12.7 mm) diameter. For 0.6 in. (15.2 mm) diameter

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strands, the transfer length tends to decrease with the increase in concrete strength. Over

a one year period, transfer lengths increased an average of 6 % with a maximum increase

of 20 %. The transfer lengths at the cut end were higher than those at the dead end by

about 20 to 30 %. It was concluded that the transfer length is proportional to the diameter

of the strand except for 0.6 in. (15.2 mm) diameter strands. The latter was referred to a

surface weathering during the transit of the 0.6 in. (15.2 mm) strands [Barnes et al, 1999].

Tabatabia and Dickson (1993), The History of the Prestressing Strand

Development Length Equation

Tabatabia and Dickson [1993] conducted this study to determine how the development

length equation in the ACI code of this time was introduced [Tabatabia and Dickson,

1993]. The development length equation was first introduced in 1963 based on the

research conducted by the Portland Cement Association (PCA) in the 1950s. A. H.

Mattock introduced the relationships for transfer length and flexural bond length

[Mattock, 1962]. The members of ACI Committee 423 modified Mattocks equation into

the development length equation existing in the ACI code [Bennett, 1963].

Ld=(1.11 fps-0.77fse)D (Eqn. 2.1)

Ld=(fps-23fse)D (Eqn. 2.2)

Equation 2.1 is introduced by Mattock to define development length. The ACI

committee modified [ACI 318, 63] Mattocks expression as shown in Equation 2.2. In

these two equations the parenthetical expression is considered dimensionless, where fps is

the ultimate stress of prestressing strand in ksi, fse is the effective prestressing stress in

ksi, and D is the nominal diameter of the strand in inches.

Mattocks expression was based on the theory of Janney that the development length

of prestressing strand can be divided into transfer length and flexural bond length. For the

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that general bond slipping occurs when the peak of flexural bond wave reaches the

prestressing transfer zone. Due to the difficulty of codifying this definition, Mattock

defined the flexural bond stress by re-appraisal of measurements of beam tests at

ultimate strength by Hanson and Kaar. He constructed a straight line relationship

between: a) the increase in strand stress beyond the effective prestressing stress at general

bond slip and at ultimate; and b) the available flexural bond length (the difference

between total length and transfer length) normalized to the strand diameter. He stated that

the line appears to be a reasonable mean line for the points representing general bond

slip, and yet is not over-conservative at large embedment lengths. Mattocks proposed

expression for flexural bond length is shown in Equation 2.5.

Ld-Lt=(fps- fse0.9) D (Eqn. 2.5)

Apparently, the members of ACI committee 423 changed Mattocks expression

for flexural bond length which was not conservative to all of the specimens to the form:

Lfb=(fps-fse) D (Eqn. 2.6)

Ld=13fseD+(fps-fse) D (Eqn. 2.7)

In Equation 2.6, Lfb is the flexural bond length in inches. Equation 2.7 defines the

development length as a summation of transfer length and the flexural bond length. This

is the same Equation (Eqn. 2.2) proposed by the ACI committee in 1963 written in a

different form.

1.1 PILESTO CAST-IN-PLACE CAPS CONNECTIONS

Harries and Petrou, (2001), Behavior of Precast, Prestressed Concrete Pile to Cast-

in-Place Pile Cap Connections

Harries and Petrou [2001] examined connections between precast prestressed piles and

cast-in-place reinforced concrete pile caps. Their objective was to determine if plain

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connections can be used instead of the connection detailing used by South Carolina

Department of Transportation (SCDOT) at this time.

The connection between pile and bent-cap should be able to develop the nominal

moment capacity of the pile. Two models were used for modeling this type of behavior.

The first model was introduced by Mattock and Gaafar [Mattock and Gaafar, 1982] while

the second model was proposed by Marakis and Mitchell [Marakis and Mitchell, 1980].

The two models deal with the embedment of steel in concrete, yet they differ in the

assumption of the stress distribution.

The experimental test consisted of two identical 18-inch (457 mm) square piles,

with a length of 18 ft (5.5 m) to check the validity of these models. Each pile was

reinforced with 8- in. (12.7 mm) diameter low relaxation strands with a prestressing

force of 31 kips (201 kN). W6 plain wire spiral was used with five turns at 1 in. (25.4

mm) pitch in the driving head. The pitch was then increased to 1.5 in. (38.1 mm) for 80

turns and 5 in. (127 mm) beyond that. The 28-day compressive strength of concrete was

found to be 6,700 psi (46.2 MPa). The caps used were identical in reinforcement using

no.7 bars on the top and bottom with no.3 ties spaced at 6 in. in the transverse direction.

Each cap had a different concrete strength, embedment length, and number of no.7 bars.

The details of the first cap were done according to SCDOT standard practice while the

details of the second cap were chosen to represent a worst case scenario.

The piles were tested as cantilever beams by orienting them in a horizontal

position. A lateral cyclic load was applied at a distance of 146 in. (3.7 m) from the face of

the cap, while a 200 kips (890 kN) constant axial force was applied to the piles to

represent a loaded bridge during a seismic event. The lateral load versus lateral

displacement hysteretic loops was developed from the recorded data. Specimen No. 1

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failed at 4 times the deflection at first yield due to crushing of concrete. Specimen No. 2

failed at 5 times the deflection at first yield due to rupture of the prestressing strands.

A pull out test was conducted on the piles due to the fact that piles may ratchet

out of the cap during cyclic axial loading in a seismic event. They did not find a

noticeable effect at a load of 75 kips (334 kN).

It was found that the two models are valid in the case of precast prestressed

concrete piles embedded in CIP caps. A parametric study was performed with different

dimensions and embedment lengths to determine the ultimate moment capacity and

nominal shear for each case. Other connection details did not significantly increase the

capacity of the connection except at extreme conditions and large deflections. Finally, the

authors recommended the embedment length to be equal the width of the pile with an

absolute minimum of 12 in. (305 mm). This embedment length is generally shorter than

the development length required for the strands to reach their nominal capacity, therefore

the full flexural capacity of the pile might not be achieved at the connection and it is the

responsibility of the designer to investigate this [Fuziol, 2007].

Shahawy and Issa, (1992), Effect of Pile Embedment on the Development Length of

Prestressing Strands

Shahawy and Issa [1992], published the findings of their research with precast

prestressed concrete piles embedded in simulated caps under lateral loading.

The testing program consisted of nineteen 14 in. square prestressed concrete pile

specimens having eight 1/2 inch (13 mm) diameter prestressing strands tested in the

primary investigation. The specimens were 12 ft (3.66 m) long cut from 80 ft (24.4 m)

long prestressed concrete piles. 5-gauge steel was used as spiral reinforcement that varied

in pitch depending on location. End sections of the original prestressed concrete piles

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were provided with more spiral reinforcement than the interior sections (middle sections).

Four embedment lengths of 36, 42, 48, and 60 in. (914, 1067, 1219, 1524 mm) were used

in the study. Cores of 6 in. (152 mm) were taken from the specimens to determine

concrete compressive strength.

The test was designed in order to simulate the behavior of a cast-in-place (CIP)

bent cap. For this purpose a test frame was used to restrain the pile against translation

and rotation. In order to determine the value of confining stress exerted from the

shrinkage of the cap an initial test was conducted to determine a realistic value for the

clamping force. This was accomplished by casting a bent cap representing an end

segment with a pile having the same dimensions of the piles used in the investigation

placed at the center. The pile was instrumented with vibrating wire strain gages along the

embedment length. After 28 days, the average confining stress was found to be 525 psi

(3.6 MPa). Using this average confining stress value as an upper limit, a clamping force

of 200 kips (888 kN) was applied to the upper and lower faces of the embedment length

of the pile to represent the confining stress. The confining stress varied with embedment

length resulting in 397, 340, 298, and 238 psi (2.74, 2.34, 2.05, and 1.64 MPa) for

embedment lengths of 36, 42, 48, and 60 in. (914, 1067, 1219, and 1524 mm),

respectively. The maximum confining stress value was taken to be 75 percent of the

measured average confining stress.

A hydraulic jack placed at 6 ft. (1.84 m) from the face of the supporting frame

was used to apply lateral load on the piles in increments of 3 kips (13.3 kN) up to a load

of 18 kips (80.1 kN), after this the increments were much smaller until failure was

achieved. At each load step, cracks were marked and displacements and strains were

recorded.

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The piles were next analyzed using a nonlinear material model. The time-

dependent effects due to load history, temperature history, creep, shrinkage, aging of

concrete, and relaxation of steel were considered in the computer program. The program

was used to calculate the structure response through the elastic and inelastic range up to

ultimate load. At each load step nonlinear equilibrium equations using the displacement

formulation of the finite element method were derived for the geometry and material

properties.

The effect of shear confinement was examined. Shear confinement was found to

have a slight effect in terms of the moment capacity of the piles. The ultimate moment of

the piles cut from the end sections was slightly higher than those of the piles cut from the

middle section by about 6%. The experimental slipping stress of the prestressing strands

was determined by measuring the strain along the length of the strand at various levels of

load till failure. For development length, the embedment length of the piles was

compared to the theoretical development length required to obtain the same slipping

stress using three different equations;

1. The current ACI code equation [ACI 318, 89] for development length of

prestressing strands which is shown in Equation 2.8. In this equation Ldis the

development length (in.), se is the effective stress of prestressing strand (psi),

ps is the nominal flexural strength of prestressing strand (psi), and db is the

nominal diameter of prestressing strand (in.).

Ld= se3000 db + ( ps- se)1000 db (Eqn. 2.8)

2. A modification to Eqn. 2.8 (current ACI approach) as proposed by Shahawy

and Issa. The proposed modification incorporates an average bond stress term

in the second part of the equation as shown in Equation 2.9, where ave is the

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calculated average bond stress (psi) which can be calculated using Equation

2.10. In this equation, P is the resisting steel strength based on the strand

slipping stress at failure (lbs.), T is the resisting concrete strength which is

assumed to be zero at ultimate due to cracking (lbs.), le is the available

embedment length (in.), and dbis the nominal strand diameter (in.).

Ld= se3000 db + ( ps- se)4 ave db (Eqn. 2.9)

ave =(P-T)( le db) (Eqn. 2.10)

3. An equation proposed by Zia and Mostafa as shown in Equation 2.11 [Zia and

Mostafa, 1977]. Their approach utilized the same parameters used in

calculating the development length of prestressing strands with the exception

of two terms, si which is the stress in prestressing steel at transfer (ksi), and

c is the compressive stress of concrete at the time of initial prestressing (ksi).

The effective stress of prestressing strand, se, and the nominal flexural

strength of prestressing strand, ps, should be used (ksi).

Ld= 1.5 sic' db-4.6+1.25( ps- se) db (Eqn. 2.11)

Comparing the different results, Shahawy and Issa indicated that the ACI code

equation is conservative when confining stress is applied to the concrete section and the

Zia and Mostafa proposed equation is even more conservative than the ACI code

equation. The Shahawy and Issa proposed equation (Eqn. 2.9) has a good match with

experimental data. However, the Shahawy and Issa equation has the notable disadvantage

that it can only be used when the slipping stress of the prestressing strand is known.

Finally, it was found that the presence of confining stress had a remarkable

beneficial effect on the development length of prestressing strands. The authors

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concluded that using ACI code equation under such conditions will lead to very

conservative values.

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CHAPTERIII

BENT-CAP CONFINING STRESS EFFECTONTHE SLIPOF PRESTRESSING

STRANDS1

1 Mohamed K. ElBatanouny, Paul Ziehl, Timothy Mays, and Juan M. Caicedo.

Submitted to ACI Journal, 8/25/2010. Status peer review.

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3.1 ABSTRACT

A lateral cyclic loading test has been completed involving four prestressed concrete pile

specimens embedded into cast in-place (CIP) reinforced concrete bent-caps. The piles

used were 18 in. (457 mm) square x 18 ft (5.5 m) long and were plainly embedded into

the bent-cap to depths of 18 in. (457 mm) and 26 in. (660 mm). The tests were conducted

to determine the moment capacity and ductility of the connections along with the slipping

stress of the strands. Analytical models were created in which the slipping stress of the

strands was calculated using ACI code equation for development length. In all cases the

connection achieved more moment than expected due to the confining stress produced

from the bent-cap which tends to decrease the development length of the strands. A

modified equation that accounts for the beneficial effect of concrete confinement due to

shrinkage is introduced.

Keywords: Embedment length, Development length, Slipping, Confining stress,

Shrinkage.

3.2 INTRODUCTION

Precast prestressed piles are commonly used in the bridge construction industry in South

Carolina and throughout the United States, often in combination with cast in-place (CIP)

reinforced concrete bent-caps. However, the performance of the connection between

prestressed piles and the CIP bent-caps under seismic loading is not particularly well

understood. The SCDOT Seismic Design Specification for Highway Bridges [SCDOT,

2008]1 is used by the SCDOT in conjunction with the AASHTO LRFD Bridge Design

Specifications [AASHTO, 2004]2 to design bridge structures to meet the demands of

seismic events.

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The current SCDOT standard connection detail utilizes a plainly embedded

prestressed pile with embedment distance of one pile diameter with an allowed

construction tolerance of 6 inches (152 mm). Plain embedment requires no special

detailing to the pile end or the embedment region and no special treatment of the pile

surface such as roughening or grooving. This connection detail is very cost effective as it

requires no special equipment or time consuming labor processes in the field. However,

the ductility and moment capacity of such connections is not well understood. The short

embedment length is often much less than the length required for developing the full

capacity of the prestressing strand. Given that South Carolina is the most seismically

active state in the eastern United States [USGS, 2009]3, this investigation has been

conducted to determine the structural performance of such connections under cyclic

loading.

It is not clear if ACI 318 Eqn. (12-4)4 for development length is applicable to

slipping stress of prestressing strands used in precast piles embedded into CIP bent-caps.

This question arises from the fact that the ACI code equation was developed for the case

of a superstructure which is different from that of piles embedded into CIP bent-caps. In

the case of vertical piles, development length of the prestressing strand is affected by

confining stress created from the shrinkage of the bent-cap. It is also mentioned that zero

moment transfer is generally assumed (pinned connection)5,6 for these connections.

ACI 318 Eqn. (12-4) may be overly conservative when used to calculate the

slipping stress of the prestressed strands used in piles, and by extension may also be

overly conservative for determining the minimum adequate length of embedment into the

bent-cap to develop the full capacity of the prestressing strands.

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In this paper, the adequacy of the ACI code equation is investigated by comparing

experimental results to moment-curvature analysis7. Experimental results of moment vs.

deflection were obtained by subjecting precast prestressed piles embedded into CIP bent-

caps to pseudo-static reverse cyclic loading. A definition of confining stress is presented

as being dependent on the amount of shrinkage experienced by the bent-cap. This

definition is used to modify ACI 318 Eqn. (12-4) by adding the beneficial effect of the

confining stress.

3.3 R ESEARCH SIGNIFICANCE

Several investigations have been conducted to examine pile to bent-cap connections and

subsequent development length of prestressing strands8,9. The investigations generally

concluded that the actual development length of prestressed strands used in piles

embedded in CIP bent-caps was less than the theoretical one. The increase in strand

stress is referred to as confining stress produced as a result of bent-cap shrinkage. In this

research, bent-caps were cast and allowed to shrink for various periods of time and an

equation for calculating the confining stress is developed in this paper. This equation is

used to modify ACI 318 Eqn. (12-4) by including the beneficial effect of these stresses

and experimental results are compared.

3.4 EXPERIMENTAL PROCEDURE

3.4.1Test specimens

Four specimens were fabricated for these tests. The specimens were created to represent

interior pile to bent-cap connections and all piles used were identical. Bent-caps were

cast around the piles with two different embedment lengths, 18 in. (457 mm) and 26 in.

(660 mm). The shorter embedment length is representative of typical SCDOT design

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procedures which require a plain embedment to a length of one pile diameter with a

construction tolerance of 6 in (152 mm). Three such piles were fabricated and tested.

For the first and second specimens, portions of the data were lost due to

equipment malfunction and therefore the results of these two specimens are combined

and referred to as specimen BC-18-1 (for Bent Cap with 18 inch embedment, first such

specimen). The remaining specimen with 18 inch embedment is referred to as BC-18-2

(for Bent Cap with 18 inch embedment, second such specimen). The longer embedment

length was simply the longest that could reasonably be obtained for typical SCDOT CIP

bent-caps with consideration to construction tolerances (specimen BC-26-1; for Bent Cap

with 26 inch embedment, first such specimen).

3.4.2Piles

All pile specimens were cast January 23rd 2009 by Florence concrete products of Sumter,

SC. Each pile had an 18 in. (457 mm) square cross section and a length of 18 ft (5.5 m).

Nine in. diameter low relaxation strands [Fu=270 ksi (1861.6 MPa)] were stressed to

31,000 lb (137.9 KN) and encased in W6 spiral wire. Piles were cast with class 5,000 psi

(34.5 MPa) concrete reaching a 28 day compressive strength (fc) of 8,278 psi (57.1

MPa). Figure 3.1 shows the layout of the piles.

3.4.3Bent-caps

Bent-caps were cast in the University of South Carolina structures laboratory according

to the current SCDOT bent-cap design with the exception of reinforcement. The

prototype design called for a bent-cap 3 feet (0.9 m) square in cross section with the piles

placed at 7 feet (2.1 m) on center. Reinforcement of the bent-cap has been marginally

reduced from standard SCDOT design in order to represent a worst case scenario. Caps

were cast with class 4,000 psi (27.6 MPa) concrete. Cap formwork was modified so that 2

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in. (51mm) PVC pipe could be fit through the cap in four separate locations to secure

caps for testing. Two of these pipes were fit through the form both at one foot (0.3 m)

from the bottom and one foot (0.3 m) from the top of the cap.

3.4.4Experimental setup and instrumentation

The four specimens were tested with the vertical axis of the pile parallel to the

Universitys laboratory floor. The specimens were held in this position by means of a

specially constructed reaction frame used in previous experimentation10. Four steel rods

ran through PVC pipes placed into the cap to the reaction frame to secure the caps in

place.

These specimens were tested with an axial compressive load of 50 kips (222 KN)

representing a dead load imposed by the bridge superstructure. The magnitude of the

axial load was determined based on a parametric study of South Carolina bridges. The

axial load was applied at the end of the pile by two hollow core hydraulic rams reacting

against a modified steel W shape. The reaction frame mentioned above was fabricated

with pins on either side of the bent-cap allowing threaded steel rods to connect these pins

to the W shape and hydraulic rams. This allowed the load to be applied such that it

remained in plane parallel to the longitudinal axis of the pile throughout a full range of

displacements, thereby minimizing P- effects. Figures 3.2(a) and 3.2(b) show a view of

the test setup.

Each specimen was instrumented with two string potentiometers to measure

displacement at a distance of 156 in. (3.96 m) from the soffit. Also, to measure curvature,

4 LVDTs displacement transducers were placed in series on each face of the pile in the

plane of displacement. Five strain gauges were welded to the longitudinal reinforcement

within the bent-cap. For specimen BC-18-2 and specimen BC-26-1 slipping was

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Testing for BC-18-2 began 113 days after casting. A 235 kip capacity (1045.3

KN) hydraulic actuator was used throughout the test for this specimen. Due to the use of

this actuator the displacement increments used in this test were less coarse than for

previous tests. The test was terminated at displacement of 8 in. (203 mm).

Specimen BC-26-1 was tested 136 days after casting of the bent-cap. The same actuator

used for specimen BC-18-2 was used for this specimen with the same pattern of

displacements; however the maximum displacement this specimen was 8 in. (203 mm).

3.1 MOMENT CURVATURE MODELS

Numerical models were developed using a moment-curvature program. For unconfined

concrete, the material properties were obtained from the report of the piles obtained from

the fabricator. Confined concrete strength depends on the amount of the confinement

produced from stirrups and this is accounted for within the program.

3.5.1Stress in prestressing strand

The stress developed in the prestressing strand is governed by the amount of the

development length available. If sufficient length is available the strand will be able to

develop its full nominal stress. ACI 318 Eqn. (12-4) for development length of

prestressed strand was used as a benchmark for this study (Equation 3.1).

Ld= se3000 db + ( ps- se)1000 db (Eqn. 3.1)

where;

Ld= development length, in.

se = Effective stress of prestressing strand, psips = Nominal flexural strength of prestressing strand, psi

db = Nominal diameter of prestressing strand, in.

For the case of piles, the available development length is equal to the embedment

length of the pile into the bent-cap. This length is less than the theoretical development

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length required to develop the full capacity of the strand, consequently this equation

predicts that the prestressed strand will slip prior to reaching its capacity. To implement

this equation, losses after one year were computed and the effective stress was calculated

as se =156 ksi (1,077 MPa). According to the grade of steel used, it was assumed that ps

= 270 ksi (1,862 MPa). Using development length values of 18 in. (457 mm) for the BC-

18-1 and BC-18-2 and 26 in. (660 mm) for BC-26-1, it was found that the predicted

strand slipping stress using this equation would be 108 ksi (745 MPa) and 156 ksi (1,076

MPa), respectively.

3.1 DEVELOPMENT LENGTHAND CONFINING STRESS

A significant difference was found between the experimental and predicted results for all

specimens as shown in Table 3.1. This difference is considered to result from the

confining stress (also referred to as clamping force) produced by the bent-cap. This

confining stress is developed due to shrinkage of the bent-cap relative to the embedded

pile. Because the piles were cast one year prior to the bent-cap, the shrinkage of the pile

is small compared to that of the bent-cap and is neglected. The current ACI code

equation (Equation 3.1) used for the modeling of the piles does not directly address the

effect of confinement on the development length of prestressed strands and therefore

modification to this Equation is proposed.

Mattock and members of the Prestressed Concrete Committee proposed the

expression for development length found in the ACI code. This equation divides the

development length of prestressing strands into two parts; transfer and flexural bond

length. The equation used the results of the study done by Kaar, LaFraugh and Mass

(1963)11 and also Hanson and Kaar (1959)12 who stated a value of average transfer bond

stress, t = 400 psi (2.76 MPa). Each of these tests was conducted on structures not

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subject to a confining stress (specimens in these test cases were limited to pre-tensioned

beam specimens or rectangular prisms). Development of the expression for the flexural

bond length had a different approach. Janney (1954)13 indicated that the general bond

slip is caused when the peak of the high bond stress wave reaches the prestress transfer

zone. This leads to a reduction in the frictional resistance due to the Hoyer effect14 and

general bond slip occurs. Also, Hanson and Kaar did not state a value for the average

flexural bond stress at the point of general bond slip, yet they agreed with the explanation

by Janney. Using data obtained from Hanson and Kaars beam tests, Mattock constructed

a straight line relationship by subtracting the estimated transfer bond length from the

embedment length of strand. The increase in strand stress due to flexure was calculated

by subtracting the effective stress due to prestressing from the value of strand stress

occurring at the load causing slip. The equations for the transfer length and flexural bond

length are shown in Equation (3.2) and Equation (3.3) respectively. While not directly

stated, it is implied that the value of flexural bond stress fb = 140 psi (0.96 MPa).

Lt= Aps*fseo *t db= fse7.36*t db= fse3000 db (Eqn. 3.2)

Lfb = fps-fse7.36*fb db= fse1000 db (Eqn. 3.3)

where;

Lt = transfer length, in.

Lfb = flexural bond length, in.

o = strand perimeter = 43**db

Aps = strand cross sectional area = 0.725* *db2/4

3.6.1Confining stress

Due to shrinkage of the bent-cap, large compressive forces starts to affect the pile causing

clamping force which is also known as confining stress. This compressive stress plays a

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part in increasing the average transfer bond stress and the average flexural bond stress,

thereby decreasing the required development length for development of the full tensile

capacity of the prestressing strand. Confining stress is dependent on several variables,

which include pile stiffness, bent-cap stiffness, dimensions of the pile/bent-cap system,

time between casting of pile and casting of bent-cap, and time between casting of the

bent-cap and loading of the specimen. With consideration to these variables, Equation 3.4

is introduced to estimate the confining stress. Due to the unsymmetrical dimensions of

the bent-cap the value of confining stress will change from one side to the other and the

use of an average value is proposed. Because shrinkage is time-dependent, the resulting

confining stress predicted with this equation varies depending on the time between

casting of the bent-cap and loading, hence confining stress will increase with increasing

time.

c = kp.kpckp+kpc.sh.lpc (Eqn. 3.4)

In Equation (3.4) c is the predicted confining stress (psi), kp is the stiffness of the

pile (lb/in.), kpc is the stiffness of the bent-cap in a particular direction (lb/in.) (generally

transverse or longitudinal direction), sh is the shrinkage strain after a certain time (in./in.)

and lpc is the distance from the edge of the bent-cap in a particular direction to the face of

the pile (in.). Equation (3.4) was developed assuming that the pile and the bent-cap work

together as springs in series subjected to the shrinkage strain of the bent-cap (neglecting

the shrinkage of the pile after approximately one year). Considering a unit area of pile

and bent-cap; kp can be calculated as Ep/lp where Ep is the Youngs modulus of concrete

used in the pile (psi) calculated from ACI 318-08 section 8.5 where [ Ep = 57000 fc ]

and lp is half the pile diameter (in.). kpc can be calculated as Epc/lpc where Epc is the

Youngs modulus of concrete (psi) used in the bent-cap and lpc is the distance from the

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value of confined transfer bond stress can be taken as the value of the average bond stress

(400 psi [2.76 MPa]) adding to it the confining stress multiplied by the friction

coefficient between steel and concrete . The confined transfer bond stress can be

calculated using Equation (3.6).

tc = 400 + . cav (Eqn. 3.6)

where;

cav = c1+c22 (Eqn. 3.7)

In Equation (3.7) cav is the average confining stress (psi) , c1 and c2 are the

confining stress in directions 1 and 2 (generally longitudinal and transverse directions)

respectively and is the coefficient of friction between steel and concrete which is taken

as described in the PCI design handbook17; =0.4.

For the confined flexural bond stress fbc (psi), the same approach was used

assuming that the confining stress will only affect the friction stress. Due to the cracks

which are formed in the average flexural bond stress zone, the Hoyer Effect is reduced

and fb is implied in the ACI code equation to be equal to 140 psi (0.96 Mpa). This will

lead to a decrease in the friction forces resulting from the confining stress. A ratio

between the average transfer bond stress and the average flexural bond stress was used to

decrease the effect of the confining stress where t/fb = 2.86. Therefore, Equation (3.8) is

introduced to find the average flexural bond stress including the effect of confining stress.

fbc = 140 + .cav2.86 (Eqn. 3.8)

3.1 PROPOSED MODIFICATIONSTO EXISTING ACI EQUATION

In order to modify the existing ACI equation (Equation 3.1), the values oftand fb are

replaced with tc and fbc respectively. In the case of confining stress being present, this

will lead to an increase in the values of average bond stress and a decrease in the

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development length. As the confining stress is proportional to time, any increase in time

will allow for further shrinkage to occur and yield a further increase in the confining

stress, thereby decreasing the development length. Therefore, it is proposed to modify the

existing ACI code to account for the beneficial effect on confining stress. Equation (3.9)

is the resulting modified equation.

Ldc = fse7.36*tc db +fps-fse7.36*fbc db (Eqn. 3.9)

The first part of the equation represents the transfer length while the second part

represents the flexural bond length, where Ldc is the confined development length (in.).

Equations (3.6) and (3.8) are used to define the values for average transfer bond stress

including confining stress and average flexural bond stress including confining stress

respectively.

3.7.1Modified Moment-Curvature models

New numerical models were developed using the moment-curvature program. The

characteristics for the unconfined and confined concrete remain the same. Equation (3.9)

was used for calculating the slipping stress of the prestressed strands.

3.1 R ESULTSAND DISCUSSION

The main experimental and theoretical results are introduced in this section. Results

related to strand slip, moment, and development length are discussed.

3.8.1Modes of failure

The shear capacity of the piles was sufficient to resist the applied load during testing.

Through visual inspection of the specimens, it was found that the strongest effect of

moment was at the junction between the pile and the bent-cap which appeared through

cracking of the pile and localized spalling of bent-cap concrete at the interface. Cracks

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were distributed along the piles in the area of the plastic hinge zone with the larger cracks

located near the interface. Due to the longer embedment length of specimen BC-26-1, the

distribution of cracks was over a length of approximately 54 in. (1,372 mm), while for the

other specimens BC-18-1 and BC-18-2 the largest distance for crack distribution from the

soffit was about 29 in. (737 mm). Further, at higher levels of load some minor cracking

occurred in the bent-cap for specimen BC-26-1 which was not the case for the specimens

with 18 inch embedment. Figures 3.6(a) and 3.6(b) show pictures of cracking in the piles.

During casting of specimens BC-18-2 and BC-26-1, the piles were rotated to

avoid top bar effect unlike specimen BC-18-1. Deterioration of moment capacity of the

piles after achieving the ultimate moment capacity of the section is due to cracking,

spalling of concrete and strand slippage. Specimens BC-18-2 and BC-26-1 were

specially instrumented with two LVDTs to measure end slipping of the top and bottom

strands, end slip results for specimen BC-26-1 are shown in Figure 3.7. The

unsymmetrical distribution of the nine strand pattern causes a difference in the slipping

results of the top and bottom strands. The maximum recorded slippage of the top and

bottom strands for specimen BC-26-1 was 0.4 in. (10 mm) and 0.26 in. (6.6 mm)

respectively. Also, for the top strand the slipping began at a smaller value of moment. It

was noticed when the specimen was failed that some of the wires forming the strands

failed, which is reasonable as the strands reached their nominal stress of 270 ksi (1,772

MPa) for this specimen. For all other specimens, the strands exhibited localized buckling

due to the high compressive forces applied.

3.8.2Effect of confinement

Confinement stress was calculated for each of the specimens. As the shrinkage of

concrete is proportional to time, giving more time for the specimen between casting and

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testing leads to higher confining stress. This leads to theoretical enhancement of the

development length. This enhancement is illustrated in Figure 3.8 which shows the

decrease in the theoretical required development length for the strand to reach its nominal

capacity. From the current ACI code equation (Equation 3.1) the strand needs about 82.9

in. (2110 mm) to reach its nominal capacity, while using Eqn. (3.9) (modified ACI

equation accounting for the beneficial effect of confining stress) for specimen BC-18-1

the strand needs only 36.7 in. (932 mm). This value is less than half of that required by

the ACI code equation. As specimen BC-18-1 incorporated data from two similar

specimens, the data of the specimen forming initial displacements was used in calculating

the theoretical required development length. For specimen BC-18-2, Eqn. (3.9) shows

that an embedment depth of 29.7 in. (754 mm) is required by the strands to reach its

nominal capacity. If Eqn. (3.9) is used for the case of BC-26-1 a distance of 31.8 in (808

mm) is needed by the strand to reach its nominal capacity. Thus, using the values of

embedment lengths for all specimens, the theoretical slipping stress can be calculated

using the ACI code equation and Eqn. (3.9). As shown in Table 3.2, the ratio (expressed

as percentage) between the theoretical slipping stress calculated from Eqn. 3.9 and the

one calculated from ACI code equation (Equation 3.1) is 171%, 189%, and 153% for

BC-18-1, BC-18-2, and BC-26-1 specimens respectively. Also, the ratio between the

predicted slipping stress calculated from Eqn. (9) and the experimental slipping is 99.5%,

127%, and 88.6% for BC-18-1, BC-18-2, and BC-26-1 specimens respectively. For

specimen BC-18-2 Eqn. (3.9) somewhat over estimates the increase of the slipping stress

by about 27%, while for the other two specimens (BC-18-1 and BC-26-1) Eqn. (3.9)

underestimates the slipping stress when compared with experimental results.

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3.8.3Moment capacity

The experimental moments were recorded in two directions (up and down), therefore the

experimental moments introduced in this discussion will be taken as the average of the

two. The moments achieved by all the specimens were very close to those calculated

from the moment-curvature program using the slipping stress calculated from Eqn. (3.9).

For BC-18-1, the moments calculated from the moment-curvature analysis using the

current ACI equation (Equation 3.1) for determining the slipping stress is about 70% of

that achieved by experimentation, while this percentage is 99.8% comparing the moments

calculated from the moment-curvature analysis using the modified ACI equation

(Equation 3.9) to the experimental moments. Specimen BC-18-2 achieved 78.9% of the

experimental moment when ACI code equation was used in calculating the slipping stress

used in the moment-curvature analysis, while using Eqn. (3.9) this percentage is 119%.

For specimen BC-26-1, the moments calculated from moment-curvature analysis using

the current ACI code equation for determining the slipping stress is about 76.3% from

that achieved experimentally, while this percentage is 97.6% comparing the moments

calculated from the moment-curvature program using Eqn. (3.9) to the achieved

experimental moments. The difference between the experimental results and the results

using Eqn. (3.9) increase when the time allowed for shrinkage increases. Due to the

longer embedment length and higher confining stress for the specimen BC-26-1 the

achieved experimental moment was equal to the calculated ultimate moment capacity of

the pile.

Theoretical results using the proposed equations showed that specimen BC-18-

2 will have higher moment capacity than specimen BC-18-1 as it experienced higher

confining stress, this did not occur experimentally as the ratio between the achieved

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experimental moment and the calculated ultimate moment capacity of the section was

84.8% for BC-18-1 and 75.2% for BC-18-2. This trend differs somewhat from the

theoretical results which predict higher moment capacity for BC-18-2 than BC-18-1.

Plots of moment vs. displacement (at 156 inches from the soffit) for specimens BC-18-1,

BC-18-2, and BC-26-1 are shown in figures 3.9, 3.10 and 3.11 respectively.

3.1 SUMMARYAND CONCLUSIONS

Efficiency of the plain connection between precast prestressed piles and CIP bent-caps

was examined using full-scale specimens that are representative of typical SCDOT

design procedures. Numerical models were created for the specimens where the slipping

stress was determined using ACI 318 Eqn. (12-4) (Equation 3.1 in this paper) for

development length of prestressed strands. Due to the presence of confining stress

produced by the shrinkage of the bent-cap onto the precast pile, the results of the model

did not match well with the experimental results. Therefore, an equation for calculating

the confining stress has been developed and is introduced. The proposed equation (Eqn.

3.9 of this paper) is a modification to the current ACI code equation where the beneficial

effect of the confining stress is taken into account in the average transfer and flexural

bond stresses. Using the modified equation, the numerical models were revised and

compared to the experimental results. The conclusions of this study can be drawn as:

1. ACI 318 Eqn. (12-4) for development length of prestressed strands is overly

conservative when used for cases where confining stress is a key parameter. This

is the case for cast-in-place bent cap bridge construction. This is not surprising as

the equation was not developed for such a case.

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2. Shrinkage of CIP bent-caps causes confining stress which enhances (reduces) the

development length of prestressed strands by increasing the average transfer bond

stress and the average flexural bond stress.

3. Shrinkage is a function of time; therefore long elapsed times prior to loading of a

pile will increase the slipping stress of strands for the same embedment length,

leading to an increase in the moment capacity of the pile.

4. Equation 3.9 is useful for the case of piles embedded into CIP bent-caps, where

variables presented in this paper can be calculated from Equations 3.6, 3.7 and

3.8. This equation can also be used for similar cases where confining stress is a

key parameter.

5. In seismic regions, plain embedment of the pile into CIP bent-caps can be used

bearing in mind that the moment capacity of the pile is dependent on embedment

length and confining stress. It is also noted that the energy dissipation mechanism

for such connections is partially due to strand slippage for more shallow

embedments.

The conclusions of this research should be verified in the field as it was based on the

laboratory testing only.

3.1 ACKNOWLEDGEMENTS

The authors wish to express their gratitude and sincere appreciation to South Carolina

Department of Transportation (SCDOT) and FHWA for financing this research work.

Special thanks to Jeff Mulliken for his input and Aaron Larosche and Shawn Sweigart for

their contributions in the experimental work. Thanks to all members of University of

South Carolina structural lab. The input of Dr. Jose Restrepo related to the experimental

program is also gratefully acknowledged.

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The opinions, findings and conclusions expressed in this paper are those of the authors

and not necessarily those of South Carolina Department of transportation.

3.2 R EFERENCES

1. SCDOT Seismic Design Specifications for Highway Bridges, Version 2.0, South

Carolina Department of Transportation, Columbia, SC, 2008.

2. AASHTO, Bridge Design Specifications, 3rd Edition, American Association of

State Highway and Transportation Officials, Washington, DC, 2004.

3. USGS, Top Earthquake States. http:// earthquake.usgs.gov regional/states/

top_states.php. January 30 2009.

4. ACI committee 318, Building Code Requirements for Structural Concrete (ACI

318-08), American Concrete Institute, Farmington Hills, MI, 2008.

5. Shahawy, M., Issa, M., and Polodna, M., Development Length of Prestressed

Concrete Piles, Report No. SSR-01-90, Florida Department of Transportation,

Tallahassee, FL, 1990.

6. Shahawy, M.A., & Issa, M., Effect of pile embedment on the development

length of prestressing strands, PCI Journal, V.37, No. 6, November-December

(1992), pp. 44-59.

7. XTRACT PROGRAM, Charles Chadwell, University of California at Berkeley,

CA, 2001.

8. Martin, L. D., and Scott, N. L.,, Development of Prestressing Strand in

Pretensioned Members, ACI journal, V.73, No. 8, August 1976, pp. 453-456.

9. Zia, P., and Mostafa, T., Development Length of Prestressing Strands, PCI

journal, V.22, No. 5, September-October 1977, pp. 54-65.

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10. Harries, K.A., & Petrou, M.F. (2001). Behavior of precast, prestressed concrete

pile to cast-in-place pile cap connections. PCI Journal, 46(4), 82-93.

11. Kaar, P. H.; La Fraugh, R. W.; and Mass, M. A., Influence of Concrete Strength

on Strand Transfer Length, Journal of the Prestressed Concrete Institute, V.8,

No.5, Oct. 1963, pp. 47-67.

12. Hanson, N. W., and Kaar, P. H., Flexural Bond Tests of Pretensioned Prestresses

Beams, ACI Journal, V.55, No.7, Jan. 1959, pp.783-803.

13. Janney, J. R., Nature of Bond in Pretensioned Prestressed Concrete, ACI

Journal, V.50, No.9, Feb. 1954, pp. 717736.

14. Hoyer, E., and E. Friedrich, Beitrag zur frage der haftspannung in

eisenbetonbauteilen (Contribution to the question of bond stress in reinforced

concrete elements), Beton und Eisen 38, (March 20), 1939.

15. ACI committee 209, Prediction of Creep, Shrinkage, and Temperature Effects in

Concrete Structures (ACI 209R-92) (Reapproved 1997), American Concrete

Institute, Farmington Hills, MI, 1992, 48 pp.

16. Stocker, M.F., and Sozen, M.A., Investigation of Prestressed Concrete for

Highway Bridges, part v: Bond Characteristics of Prestressing Strand, Bulletin

503, Urbana, University of Illinois Engineering Experiment Station, 1970.

17. PCI Design Handbook: Precast and Prestressed Concrete, 6th Edition, MNL-120-

4, Precast/Prestressed Concrete Institute, Chicago, IL, 2004.

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Table 3.1: Maximum moment and slipping stress (measured and current ACI Equation)

Specimen Embedmentdepth

in. (mm)

Experimental results ACI Eqn. 12-4

Max. Momentkip-in (kN-m)

Slippingstress*

ksi (MPa)

Max. Momentkip-in (kN-m)

Slipping stressksi (MPa)

BC-18-1 18(457)

2345(247)

185 ksi(1276)

1,642(186)

108(745)

BC-18-2 18(457)

2080(219)

160(1103)

1642(186)

108(745)

BC-26-1 26(660)

2765(312)

270(1772)

2102(238)

156(1076)

* = Experimental slipping stress was determined using the value of maximum moment achievedby the specimen into the moment-curvature program.

Table 3.2: Comparison of calculated slipping stress (current ACI Eqn. and proposedEqn.)

Specimen

Epc

psi (GPa)

cav

psi (MPa)

ACI Eqn.12-4Modified ACI Eqn.

(Eqn. 3.9)

psksi (MPa)

tc,psi

(MPa)

fbcpsi

(MPa)

psksi

(MPa)

BC-18-14.21e6

(29)

1230

(8.46)

108

(745)

891

(6.14)

312 psi

(2.15 MPa)

184

(1270)

BC-18-23.99e6

(27.5)

1750

(12.1)

108

(745)

1102

(7.59)

385 psi

(2.66 MPa)

204

(1404)

BC-26-13.72e6

(25.6)

1570

(10.8)

156

(1076)

1028

(7.09)

360 psi

(2.48 MPa)

239.2

(1649)

Table 3.3: Comparison of slipping stress and maximum moment

SpecimenEmbeddepth

in. (mm)

Slipping stress Max. MomentExper.results

ksi (MPa)

ACIksi (MPa)

Eqn. 3.9ksi (MPa)

Exper.results

k-in. (kN-m)

ACIk-in. (kN-m)

Eqn. 3.9k-in. (kN-m)

BC-18-1 18 (457) 185 (1276) 108 (745) 184 (1269) 2345 (265) 1642 (186) 2340 (264)

BC-18-2 18 (457) 160 (1103) 108 (745) 204 (1407) 2080 (235) 1642 (186) 2483 (281)BC-26-1 26 (660) 270 (1862) 156 (1076) 239 (1648) 2765 (312) 2102 (237) 2699 (305)

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Metric (SI) conversion factors: 1 in. = 25.4 mm; 1 ksi = 6.895 MPa; 1 lbs = 4.45 N.

Figure 3.1: Prestressed pile reinforcing details and strand arrangement.

Figure 3.2(a): Overview for the test setup with hydraulic

actuator.

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Figure 3.2(b): Setup for loading the pileusing hydraulic cylinders.

Figure 3.3: Closeup of the LVDTs (encased in steeltubes) for measuring strand slip.

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Metric (SI) conversion factors: 1 in. = 25.4 mm; 1 kip = 4.45 kN.

Figure 3.4: Schematic of the test setup.

Figure 3.5: Pile/bent-cap confining mechanism.

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Figure 3.6(a): Crack and spalling on top face of pile withstrand visible.

Figure 3.6(b): Cracking away from interface and spallingof bent-cap.

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Metric (SI) conversion factors: 1 in. = 25.4 mm; 1 kip-in. = 0.11298 KN-m.

Figure 3.7: Strand end slippage vs. moment for BC-26-1.

Metric (SI) conversion factors: 1 in. = 25.4 mm; 1 ksi = 6.895 MPa.

Figure 3.8: Effect of confinement on development length.

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Figure 3.9: Moment vs. displacement for BC-18-1.



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CHAPTERIV

DETERMINING SLIPPING STRESSOF PRESTRESSING STRANDSIN CONFINED

SECTIONS2

2 Mohamed K. ElBatanouny, and Paul Ziehl. Submitted to ACI Journal, 10/26/2010.

Status peer review.

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1.1 ABSTRACT

Development length and slipping stress of prestressing strands subjected to confining

stress is not well quantified and the validity of the ACI code equation for development

length of prestressing strands under such conditions can be questioned. In 1992, a test

was performed on nineteen 14 in. (356 mm) square prestressed concrete piles with a

clamping force applied to these piles during testing under lateral load. The findings can

be summarized in the inadequacy of the ACI code equation for development length of

prestressing strands when used for sections subjected to confining stress. A modified

equation that accounts for the beneficial effect of concrete confinement is proposed and

compared to the published 1992 results and to the ACI code equation. Ultimate moment

capacity for sections is also compared using moment-curvature analysis by setting three

different slipping values from experimental results, ACI code equation, and the proposed

modified ACI code equation.

Keywords: Development length, Slipping, Moment capacity, Confining stress.

1.2 INTRODUCTION

The use of precast prestressed piles in bridge construction is common in the United

States; however the performance of such units under seismic loading is not well

understood. The behavior of the connection between prestressed piles and cast-in-place

reinforced concrete bent caps is particularly not well understood. Current standard

SCDOT1 connection details require the plain embedment of the pile into the bent cap one

pile diameter with a construction tolerance of 6 inches (152 mm). Plain embedment

requires no special detailing to the pile end or the embedment region and no special

treatment of the pile surface such as roughening or grooving. The ductility and momentcapacity of such connections is of concern because this short embedment length is often

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much less than the length required for development of the full tensile strength of the

prestressing strands within the embedded region.

Generally slipping stress and development length of prestressing strands is calculated

from ACI 318 Eqn. (12-4)2, yet the validity of this equation for strands in confined

sections is not proven. This question arises because this equation was introduced for the

case of superstructure elements not subjected to confining stress. However, a pile

embedded into a bent-cap is subjected to the shrinkage of the confining concrete in the

bent-cap which creates confining stress (also known as clamping force) on the pile

which serves to enhance the bond between the prestressing strand and the surrounding

concrete leading to

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