APPLICATION OF STRESS-WAVE THEORY TO PILES
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PROCEEDINGS OF THE SIXTH INTERNATIONAL CONFERENCE ON THE APPLICATION OF STRESS-WAVE THEORY TO PILES/SAO PAULO/BRAZIL/ 11 - 13 SEPTEMBER 2000
ss Edited by
Sussumu Niyama Institute for Technological Research - IP2: SLio Paulo, Brazil
Jorge Beim PDI Engenharia, Rio de Janeiro, Brazil
A.A. BALKEMA/ROTTERDAM/BROOKFl[ELD/2000
This conference was organized by the Brazilian Society for Soil Mechanics and Geotechnical Engineering
The texts of the various papers in this volume were set individually by typists under the supervision of each of the authors concerned.
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Application of Stress-Wave Theory to Piles, Niyama & Beim (eds) 02000 Balkema, Rotterdam, ISBN 90 5809 150 3
Table of contents
Preface
XIJI
Organization
xv
Acknowledgements
XVII
1 Wave mechanics and its application to pile analysis Keynote lecture: Some wave mechanics applications G.G.Goble
3
Analysis of bearing capacity of rock-socketed piles based on wave equation theory L.B. Cai
11
Pile acceptance based on combined CAPWAP analyses R. F: Stevens
17
Theoretical study on effect of pile shaft resistance on rebound during pile driving R.P Chen & Y M. Chen
29
Time effect in determining pile capacity by dynamic methods M. R. Svinkin
35
Set-up considerations in wave equation analysis of pile driving CWCho, M.WLee & M.F:Randolph
41
Drivability and performance of model piles driven into cemented calcareous sand D. Bruno, M. F: Randolph, C WCho & H A Joer
47
Automatic signal matching with CAPWAP E Rausche, L. Liang & B. Robinson
53
Combining static pile design and dynamic installation analysis in GRLWEAP E Rausche, B. Robinson & J. Seidel
59
Rules of thumb for field and construction engineer in relation to impact pile driving W J Lucieer
65
V
2 Driving equipment and recent developments -New technologies for quality assurance of piles Keynote lecture: Pile driving equipment: Capabilities and properties F: Rausche
75
Hammer system design using wave equation analysis for testing cast-in-situ concrete piles M A Mukaddarn, WM.Iskandarani & M. Hussein
91
Analysis of pile load transfer using optical fiber sensor J.-H.Oh, W-J.Lee, S-B.Lee & W-J.Lee
99
Set-up effect of cohesive soils in pile capacity M.R.Svinkin & R.Skov
107
Modulus of elasticity and stiffness of composite hammer cushions M. R. Svinkin
113
A new technique to drive piles: Down-the-holepiling A. Benarnar
117
Simplified neural network models for estimating soil resistance using dynamic pile test R. Liang & N 0.Nawari
121
Pile capacity prediction using neural networks technique A. S Dyrninsky, E. Parente-Ribeiro, C.Rornanel & J. W Beim
127
Hammer and pile cushion optimalisation G.Jonker & R.J.van Foeken
135
3 Pile integrity and low strain dynamic testing Keynote lecture: 0resundlink, grouting work under West Pylon: Integrity test of a giant pile J. Rome11
145
A study for the shape of pile with reflect-wave method Yang Wu
153
Evaluation of the performance of an existing foundation via PIT tests S B. Foa, J. H. l? Pereira, R. P Cunha & J. Carnapum de Carvalho
157
Sonic integrity test of piles-integrity effected by basement excavation in Bangkok soft clay NThasnanipan, A.WMaung, PTanseng & Z.ZAye
163
Non-destructive integrity testing on piles founded in Bangkok subsoil N. Thasnanipan, A. WMaung, T.Navaneethan & Z.Z.Aye
171
Estimation of pile head stiffness using sonic integrity testing K. Irnada, T. Matsurnoto & Y:Nakata
179
Integrity testing of cast-in-situ concrete piles associated with the construction of New Haccho Bridge X Michi, T.Matsurnoto & K Matsuda
I87
Presentation of low strain integrity testing in the time-frequency domain J. R Seidel
193
VI
First experiences in the application of the stress wave theory to foundations in Uruguay A.Gutikrrez, L.Abreu, Ch.H o f i a n n & D. Hasard
20 1
Detection and prevention of anomalies for augercast piling G. Likins, G. Piscsalko, E;: Rausche & C.M. Morgano
205
Recent advances and proper use of PDI low strain pile integrity testing G. Likins & E;: Rausche
21 1
Stress wave propagation velocity at early ages C Restrepo
219
Examination of a new cross-hole sonic logging system for integrity testing of drilled shafts S.G. Paikowsky, L. R. Chernauskas,L.J.Hart, C D. Ealy & A. I;:DiMillio
223
4 Pile-soil interaction Keynote lecture: Identification of soil-pile model interaction parameters from recorded time-displacement signals A. I. Husein Malkawi & I. M.Ayasrah
233
Load transfer analysis from increasing energy dynamic load tests in concrete piles driven in very soft clay formation J.Balech & NAoki
241
Dynamic pile testing and finite element calculations for the bearing capacity of a quay wall foundation - Container terminal Altenwerder, Port of Hamburg F: Kirsch, B. PlaPrnann, ir:Huch & W Rodatz
249
The shaft dynamic response of a pile in clay: Induced pore pressure
255
A. Benamar
An investigationof pile diameter influence in the bearing capacity on Dynamic Load Test (DLT) G I? Bernardes, C S Andreo & C Gongalves
26 1
Pile set-up in sands J. I? Seidel & M. Kalinowski
267
Modeling pore pressure generation during dynamic testing of deep foundations
275
l? L. Pinto
5 Codes Keynote lecture: The performance of the dynamic methods, their controllingparameters and deep foundation specifications S G. Paikowsky & K. L. Stenersen
28 1
LRFD design codes for pile foundations - A review G.G.Goble
305
Pile testing competitions - A critical review J. M.Arnir & B. H. Fellenius
313
VI I
The need for quality assurance in the dynamic pile testing industry J. 2? Seidel
319
High strain dynamic pile testing, equipment and practice G.Likins, E Rausche & G.G.Goble
327
Dynamic load testing and Statnamic load testing for acceptance and design of driven piles in Japan Matsumoto, K Fujita, 0.Kusakabe, M. Okahara,N Kawubata & S. Nishimura
335
Ethics and money. Are they compatible? H. Goldemberg & J.J. Goldemberg
345
6 High strain dynamic testing of driven and cast in situ piles - Dynamic testing of large piles Keynote lecture: Pile acceptance criteria for large diameter and cast in situ piles R. E Stevens
35 1
Analyzing the bearing capacity mechanism of large diameter diving casing cast-in-situ concrete piles by using high strain dynamic testing Xi Liang
369
Analysis of dynamic load tests on steel rails piles EM.A. Lima, J.C.A.Cintra & NAoki
375
A discussion of penetration matching on high strain dynamic pile testing D.Xu, S. Wu & L.Xiao
383
Experience gained and difficultiesin performing dynamic load test in composite piles made with steel rails G.P Bernardes, C.S.Andreo & C.GonCalves
389
The application of high strain dynamic pile testing to screwed steel piles J.G.Cannon
393
Case study on the application of high strain dynamic pile testing to non-uniform bored piles J.G.Cannon
399
High capacity dynamic load tests for bored piles in Sydney shale D.J. Klingberg & fl Mackenzie
403
Predicting uplift deflection from dynamic pile testing WG.Chambers & D.J.Klingberg
407
Applicability of dynamic load test on a toe improved pile J. Sakimoto,N Kitu, S. Nishimura & TTakeda
41 1
Dynamic pile testing practice in Finland
415
H,Jokiniemi, J. Hartikainen & I? Korkeakoski
Driving behavior of large diameter steel pipe piles 1-A? Lee & J. -SLee
Vlll
42 1
Dynamic load test of cast in place pile using a free fall hammer S. Niyama, G. Cde Campos, S. Navajas, S. C.Paraiso, C M. C.Costa & G.E. Barbosa
429
Dynamic testing of large auger pile using free fall loading system in a harbour work S. Niyama, S.Navajas & G.C.de Campos
435
Dynamic load test on high capacity pile socketed in basaltic rock S. C Paraiso, C M. C.Costa & E. Pinto Soares
441
7 SPT measurements and special field monitoring test Keynote lecture: Frequency characteristics of stress wave and penetration during SF'T K. Fujita
45 1
The application of energy conservationHamilton's principle to the determination of energy efficiency in SPT tests N.Aoki & J. CA.Cintra
457
Correlative study of Smith damping coefficient and SPT blow count R.1 Liang
461
Stress wave theory application to standard penetration test in Japan K. Fujita & M. Ohno
469
8 Vibratorypile driving - Vibration in pile driving Keynote lecture: Vibratory driving analysis A. E. Holeyman
479
Computation of ground waves due to piling C.L. Ramshaw,A. R. Selby & l? Bettess
495
The effect of pile impedance on energy transfer to pile and ground vibrations M. R. Svinkin,B. C.Roth & IT?R. Hannen
503
Determination of modulus of subgrade reaction in a pile with a vibrating apparatus M. Hilmi Acar
51 1
Analysis of crosswise vibration of pile driving R.l?Chen, B.Zhu & YM.Chen
517
Monitoring and control of dynamic effects of pile installation prior to pile driving M. R. Svinkin
525
Full-scale field-test study of dynamic soil resistance of vibratory driven sheet piles K. Viking
533
Accelerations of a driven pile and the surrounding soil E.L. Hajduk, S.G. Paikowsb, I! Holscher & l? B.J. Barends
541
9 Statnamic and other similar techniques Keynote lecture: Statnamic, the engineering of art R Middendorp IX
55 1
Keynote lecture: Three-dimensional finite element analysis of statnamic load test I:Boonyatee~M. Kimura & F: Zhung
563
Lateral statnamic load testing of model piles M.~imura& ~ ~ o o n y a t e e
569
Application of the Stress Wave method to a~tomaticsignal matching and to statnamic predictions G. Esposito, WMG.Courage & R.J.van Foeken
575
A comparative study of static, dynamic and statnamic load tests of steel pipe piles driven in sand A, ~hibata? N Kawa~ata,~ ~ a k i yY:aYoshizawa, , M.~ u ~ a s&h 1:i Ma~s~moto
583
Case studies of statnamic load testing in Japan S. Nishimura, 1:Matsumoto, 0.Kusakabe, K Nishiumi &: YYoshizawu
591
Statnamic and dynamic load tests for large diameter steel pipe piles supported by a thin bearing layer at Nagoya port in Japan Y:~ikuchi,S Nishimura & M. ~atsuta
599
Statnamic load testing using water as reaction mass M. D.~ustason? ~ * ~ ~ a I!n~ei ds d, e n d ~ & r pAG. ~ u l l i n s
609
Introducing statnamic load testing in Europe: Case studies in the Netherlands G.JJ,van Ginneken & I! Middendorp
617
The advantages and disadvantages of dynamic load testing and statnamic load testing P Middendorp, G.J.J. van Ginneken & R.1 van Foeken
625
10 Case histo~ies, pile set-^^ and c ~ ~ ~ e l a t i~o n se ~teste ethe hods ~ - P ~ e ~ i c ~~i oen~ i a ~ i ~ i ~ Keynote lecture: Improving the reliability of pile bearing capacity prediction by the dynamic increasing energy test (DIET) ~ ~ o k i
635
Correlation analyses of dynamic and static loading tests for nine piles Y:M. Zheng, J , M. Zheng & B. Chen
65 1
~ a c k - ~ ~ yofs steel e s pile driving records for quality assurance B. R. Danziger & J. S. Ferreiru
657
Evaluation of pile set-up from penetration per blow G.Axelsson & S Hintze
665
Comparative analysis of dynamic and static test of foundation pile GZhou &LWu
673
Dynamic load test and elastic rebound analysis for estimation of the bearing capacity of piles in residual soil RL R.de Albuquerque & D.de Carvalho
677
Strain dynamic testing on pressure-grouted piles Liu Xi-An & Zhang Yao-Nian
683 X
Assessment of the interface between dynamic and rapid loading tests M. B. Karkee, X Sugimura & 7:Horiguchi
689
Dynamic load testing on 102 steel pipe piles for bridge foundations on mudstone M. Hayashi, 7:Matsumoto& M. Suzuki
697
Behavior of short CFA piles in an overconsolidated clay based on static and dynamic load tests A. C.M. Kormann, R R. Chamecki,L. Russo Neto, L.Antoniutti Net0 & G.19Bernardes
707
Static and dynamic testing of the ‘Campile’ - A displacement, cast-in-situ pile D.J.Klingberg & r! Mackenzie
715
Is DLT the final word? Correlation between DLT and SLT
719
H.Goldemberg & J.J.Goldemberg Results of an international pile dynamic testing prediction event A. Holeyman, J. Maertens, N. Huybrechts & C Legrand
725
Preparation of an international pile dynamic testing prediction event A Holeyman, J. Maertens, h? Huybrechts & C Legrand
733
Case studies of dynamic load testing in Japan X Wakiya,K. Nishiumi, M. Hayashi,A. Shibata, S.Nishimura & TMatsumoto
741
Case studies of high capacity CFA pile testing in Australia S. Baycan
75 1
11 Supplement Keynote lecture: Retrospective of Sonic Integrety Tests - Its application to the control quality 757 on piles J. J. Goldemberg Author index
77 1
XI
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Applicationof Stress-Wave Theory to Piles, Niyama & Beim (eds)02000 Balkema, Rotterdam, ISBN 90 5809 150 3
Preface
The application of dynamic testing based on the wave equation theory was first introduced in Brazil in 1981. Differently from other countries, its first application was on offshore piling. At that time, a great number of jacket type platforms began to be installed in the country by the Brazilian Oil Company, in waters with depths varying from 50 to 250 meters. The use of this technique also on land job sites experimented a continuous growth since then. In 1989, low strain integrity testing was introduced. In 1994, a code on the dynamic load testing of piles was issued by the Brazilian Technical Standards Association - ABNT A couple of years later, in 1996, the same ABNT edited a revision of the Standard for Design and Execution of Foundations, which now accepts the use of dynamic load testing as one of the alternatives for bearing capacity evaluation. In the case of driven piles, it also allows for a reduction of the safety factor, from the usual value of 2 down to 1.6, provided that dynamic testing is performed on at least 3% of the piles on the job site, and that it is applied since the beginning of the piling. The present event in the Southern Hemisphere, and particularly in South America, certainly will contribute to increase the utilization of this technique, thus helping improve the quality of pile installation in this region. The support of the Brazilian Society for Soil Mechanics and Geotechnical Engineering - ABMS in hosting this conference had the purpose of trying to make the geotechnical and foundation engineering communities and the wave equation theory users work more closely together. The conference structure was modified, eliminating the traditional presentation of the papers, thus allowing more time for discussions. After five conferences, the first one organized by Dr H. Bredenberg of the Swedish Geotechnical Institute in 1984, we hope that discussions based on the papers and lectures presented in this proceedings might have contributed to improve the understanding and bettern interpretation of dynamic tests, not only for the executioners but specially for the users of the the tests.
Sussumu Niyama Jorge Beim Editors
Xlll
SPONSORSHIP
Brazilian Society for Soil Mechanics and Geotechnical Engineering SUPPORT FINEP
XIV
Applicationof Stress-WaveTheory to Piles, Niyama & Beim (eds)02000 Balkema, Rotterdam, ISBN 90 5809 150 3
Organization
ORGANIZING COMMITTEE Niyama, Sussumu Chairman Beim, Jorge Secretary Andreo, Cristiana Aoki, Nelson Bernardes, George Campos, Gisleine Coelho Cintra, Jos6 Carlos Danziger, Bernadete Falconi, Frederico E
Gonplves, Cliiudio Mello, Jayme Ricardo de Kormann, Alessander Morales Merighi, Valmir Navajas, S6rgio Paraiso, S6rgio C. Valverde, Skrgio
CONFERENCE ADVISORY COMMITTEE Lopes, Francisco Rezende Velloso, Dirceu de A. D6court, Luciano Golombek, Sigmundo Mello, Victor E B. de INTERNATIONAL,ADVISORY COMMITTEE Altaee, Ameir - Canada Amir, Joram M. - Israel Burbano, Germiin J. - Spain Fellenius, Bengt H. - Canada Fujita, Keiichi -Japan Goble, George G. - USA Goldemberg, Juan J. - Argentina Gravare, Carl-John - Sweden Gutierrez, Alvaro - Uruguay Holeyman, Alain E. - USA Holloway, D. Michael - USA Hussein, Mohamad H. - USA Kusakabe, Osamu - Japan Liang, Xu Ding - China
Likins, Garland - USA Matsumoto, Tatsunori -Japan Mendiguren, Eugenio - Argentina Middendorp, Peter - Netherlands Paikowsky, Samuel G. - USA Pinto, Paul0 - Portugal Randolph, Mark - Australia Rausche, Frank - USA Sakai, Tomoaki - Japan Seidel, Julian F! - Australia Towsend, Frank C. - USA Van Impe, William - Belgium Van Weele, A. E - Netherlands Veiga, Jose - Chile xv
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Application of Sfress-Wave Theory to Piles, Niyama L? Beim (eds) 02000 Balkema, Rotterdam, ISBN 90 5809 150 3
Acknowledgements
The editors are grateful to the following persons who helped to review the manuscripts and thus greatly assisted in improving the overall technical standard and presentation of the papers in these proceedings: Kormann, Alessander Morales Danziger, Bernadete Gravare, Carl-John Velloso, Dirceu de A. Lopes, Francisco Rezende Goldemberg, Hermh Amir, Joram M. Seidel, Julian I? Svmkin, Mark R. Randolph, Mark Kusakabe, Osamu Middendorp, Peter Goble, George G. Holeyman, Alain E. Paikowsky, Samuel G. Paraiso, Skrgio C. Matsumoto, Tatsunori L W s , Garland
XVll
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1 Wave mechanics and its application to pile analysis
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Application of Stress-Wave Theory to Piles, Niyama & Beim (eds) 02000 Balkema, Rotterdam, ISBN 90 5809 150 3
Keynote lecture: Some wave mechanics applications George G.Goble George G.Goble Consulting Engineer, University of Colorado, Boulder, Colo., USA
ABSTRACT: The solution to the one dimensional wave equation is used to determine the maximum impact force at the top of a concrete pile for a given ram weight, cushion stiffness, and pile impedance. The result is used to generate a set of curves that can be used to assist in the selection of a pile cushion as limited by maximum compression stress. A second problem, the wave equation modeling of the Statnamic load testing system, was also studied. The testing process was modeled using a wave equation computer program. A comparison of the commonly used damping model, Smith damping, with the Rausche model can be used to evaluate the appropriateness of the two models. This study indicates the desirability of the Rausche model and the necessity for fbrther study to recommend appropriate values for the Rausche damping constant. plays now generally used for presenting dynamic measurements, for many of the usefbl concepts for visually evaluating those measurements, and for current approaches to integrity testing of deep foundations. Since the basic work is complete and many direct applications of the theory well established it is unlikely that fbrther hndamental developments will occur. It is appropriate to note that, while the use of the closed form solution has been very important in developing conceptual understanding of the mechanics of pile driving, it has been less usefil for obtaining quantitative results.
1 INTRODUCTION
The interest in one dimensional wave propagation theory for understanding pile driving is more than a century old. Perhaps, the pile driving problem was a major motivation in deriving and solving the one dimensional wave equation since those people that were active in that effort were mostly Civil Engineers. The important work done in this area that relates directly to current applications to pile driving started shortly aRer the Second World War and continued to the present.
Probably the most remarkable work in developing quantitative solutions was done by E. A. L. Smith, then Chief Engineer of the Raymond Company, at that time, the largest pile driving company in the world. Before 1950, he published notes on the development of a numerical solution to the wave propagation problem applied to pile driving (Smith, 1950). Shortly aRer, a proprietary program was operating on an electronic digital computer. Smith called this program the “Wave Equation” and it was widely publicized (for example; Smith, 1957; Smith, 1960). This program may have been the first application of electronic digital computers to a civilian engineering problem.
Today it can be said, with confidence, that the use of one dimensional wave propagation theory is generally accepted and widely used in the pile driving industry. The hallmark of these applications is that they have centered on usefil applications of the theory. In this paper, two topics will be discussed, one dealing with the closed form solution of the wave equation and the other, the discrete solution. Useh1 results will be presented in each case. The one dimensional wave equation was derived and solved in the nineteenth century. Probably the most important contributions to the use of the closed form solution were made by Professor Fischer of Upsula University beginning in the early 1950’s. He developed many useh1 applications of the solution including several graphical approaches. A summary of this work was presented at the Second Conference on the Application of Stress Wave Theory to Piles (Fischer 1984). This work laid the basis for the dis-
The implementation of wave equation analysis has been painstakingly slow. A public domain program was developed in the early 1960’s at Texas A&M University and it was widely used in the offshore industry (Samson, et a1 1963). Today, the
3
GRLWEAPTMprogram (Goble & Rausche 1976, GRLWEAPTM,1998) is widely used in the United States and that usage is increasing rapidly. It is becoming standard practice to perform driveability analyses both during the foundation design stage and at the beginning of construction. However, outside the United States usage is much less common.
i3x
c dt
where A is the cross sectional area of the pile, E is its elastic modulus, c is the velocity of wave propagation, o is the stress at the top of the pile, e is the strain at the top of the pile, x defines the distance from the top of the pile to some point on the pile, and t is time. This expression is substituted into the expression for the driving force, F, in Equation (1). It is then written in terms of the displacement of the pile top and the resulting expression is rearranged and expressed in terms of the pile top velocity ,v.
2 PILE CUSHION SELECTION
A direct application 'of a closed form solution to the problem of selection of a pile cushion for driving a concrete pile will be presented. This solution can be used with the aid of graphs and it may be found to be usefiil in practice. Clough and Penzien (1975) solved the problem shown in Figure 1. The pile driving system is modeled by three elements, a rigid ram, a linear spring (cushion) and an elastic pile. The solution provides the maximum delivered impact velocity induced at the pile top during the first wave passage (among other things). This solution is used to generate curves for selection of the cushion stiffness required to limit the induced compression stress to some selected value.
k k v+-++-v Z m
=0
(4)
In Equation (4), the weight of the ram has been neglected and Z is the pile impedance, ENC. This familiar equation form has the solution v=
(A sin o t + B coso t)
(5)
The variables k/m and k/Z in Equation (4) have been replaced by a2and 205, respectively, and
If the initial conditions of the pile top velocity,v, and the ram impact velocity, vh, are imposed, the expression for the velocity at the top of the pile becomes
Figure 1. Pile Driving System Model
The equation of motion for the ram in contact with the cushion can be written F=mg-muh
(1)
where the variables are defined in Fig. 1. Continuity of displacements requires that the ram motion equal the pile top motion plus the spring deformation.
This expression is differentiated twice and substituted into Equation (1). The resulting expression is solved for F and then expressed in terms of the stress in the pile as given by one dimensional wave mechanics.
Figure 2. Values v/vh for pile stiffness, ram mass, with given impedance
4
Maximum values of v/vh were determined for a range of values of ram mass, cushion stiffness, and pile impedance.
used quantitatively so contours for four values of Z are given in Figures 3 through 6.
In order to use the curves, trial cushion stiffness and ram mass is selected, and they are divided by the pile area. Likewise, the pile impedance divided by the pile area is also determined. Using the appropriate figure, the pile velocity-ram velocity ratio is determined from the curve. This quantity is multiplied by the anticipated ram impact velocity to obtain the maximum particle velocity transmitted to the pile top. The particle velocity times the pile impedance gives the maximum impact force transmitted to the pile.
A parameter study was made for a 300 mm square pile with a full range of practical values of ram mass, cushion thickness, and concrete modulus. Values of the ratio of the pile top velocity to the ram impact velocity, dvh, were determined for a range of values of ram mass and cushion stiffness, and four values of pile impedance. An.example of the results is shown in Figure 2 where the surface of v/vh is given for a range of values of k and m with a specific value of pile impedance, 2. Such a representation cannot be
5
Figure 7. Wave equation model
It should be noted that the model does not include the hammer cushion or the helmet mass. Furthermore, the result has not been tested against field experience and such testing must be done to assure the usehlness of the results.
There can be little hope of ever accurately knowing the actual hammer efficiency prior to going to the field, Even beginning driving the hammer eficiency will only be known when measurements are available. The actual quake that exists during pile installation is dependent on the size of the pile cross section and on ~01.1properties. Quake values can be assigned based on the pile cross section but when a large quake is due to soil properties it cannot be predicted in the current state-of-the-art. Fortunately, this characteristic occurs infrequently. Experience with dynamic measurements and signal matching analysis indicates that this problem usually occurs during extended driving and is probably pore pressure dependent. It disappears in re-strike testing. A solution to this problem will depend on a better understanding of soil behavior.
3 WAVE EQUATION DAMPING MODEL Wave equation analysis is now used routinely in land pile driving practice in the United States and also in controlling the installation of piles for offshore petroleum recovery platforms. The model used for wave equation analysis is shown in Figure 7. The three principal problems that limit the accuracy of wave equation analysis are a lack of knowledge of the driving efficiency for a particular pile driving hammer, an occasional, surprisingly large quake value, q, and a lack of accuracy in the soil damping constant, j.
6
evaluated as dynamic but it is slower than the traditional dynamic test.
Damping constants have been selected based on tradition using the original recommendations of Smith (1960). They are selected based on soil type and they show a wide range of accuracy based on signal matching analyses (Rausche et al 1994). This variability can strongly affect the results of wave equation analysis. Rausche et a1 (1994) suggested a different model that is based on the work of Coyle and Gibson (1970). Coyle’s research, fbrther proven by Herema (1979), indicates that the dynamic resistance has a strongly nonlinear relationship to the velocity of pile motion. The study by Rausche et a1 indicated that the direct application of the Coyle research caused numerical problems. A suggestion by Rausche avoided those problems.
The Statnamic test was modeled using a mass of the typically used size as a “ram” in GRLWEAPTM. It was dropped on the top of the pile with a cushion having a very soft spring. The spring constant was selected by trial-and-error so that a force pulse of about 100 ms was generated. The magnitude of the peak force was controlled by the ram drop height. The test pile used in the study was a closed-end steel pipe with a length of 24 meters and an ultimate capacity of 4000 kN. An example of the force-time and velocity-time record for one of the tests is shown in Figure 8.
The traditionally used Smith model states R, = R,(1+ j,v) where Rt is the total soil resistance, R, is the current value of the static resistance, and j, is the Smith damping constant. This model becomes viscous when the static resistance is equal the ultimate resistance. Based on laboratory testing Coyle showed that a more appropriate law would be R , = R,(1+ j g v N )
(9)
where j, is the Gibson damping constant and N is an exponent, typically less than 1.0. The difficulty with this representation is that numerical problems arise when the velocity changes sign. Rausche suggested that damping be represented by
Figure 8. Statnamic Test Time
-
Force and Displacement vs.
Four cases were studied, each with a range of applied forces. They are (a) Smith damping for sand, (b) Rausche damping for sand, (c) Smith damping for clay, and (d) Rausche damping for clay. The results are given in Figures 9 through 12.
v R, = R,[1+ j R v r -v, R,
In Figure 9 the force-displacement results for the case of the pipe pile in sand is shown. The wave equation results include the top force in the pile as a hnction of time together with the pile top displacement. The maximum applied force was 5200 kN. In this case a maximum displacement of 55 mm was achieved. The method that has been recommended for use in determining the static capacity was to use the force at the time of zero velocity and adjust this force by the mass of the pile times the acceleration at that same time. This method was first suggested by Nara (1970) and was used by the Case Research project (Goble & Rausche, 1970). It can be seen in Figure 9 that this method gives quite good results when compared with the specified static capacity of 4000 Kn. The applied forces were 5200 kN, 4840
where j, is the Rausche damping constant, vx is the maximum velocity achieved up to a particular time during the blow, and R, is the maximum static resistance actuated prior to the time under consideration. Both the Smith and the Rausche model are available in GRLWEAPm. The Statnamic method (Janes et a1 1994) uses a charge of slow burning explosive to generate a force between a large mass and the pile top. As the mass is accelerated upward a downward acting force is induced against the pile top. This force has a slower rate of increase than the typical ram impact used in dynamic testing. Force pulses of the length of 80100 ms are common. The resulting test must still be
7
in a clay soil. The usual Smith damping constant result is given in Figure 11. The predicted static capacity for all three Statnamic load levels are all quite close to the known static capacity.
kn, and 4180 kN. In all three cases the capacity at maximum displacement was about 4000 kN. The adjustment for the acceleration at zero velocity (inertia force) would be small due to the low mass of the pile.
Figure 12 shows the result for Rausche damping in clay for applied forces of 8000 kN, 7150 kN, 6070 kN, and 5450 kN. The associated predicted static capacities are 6400 kN, 6400 kN, 5830 kN, and 5240 kN, respectively. All of these values are quite high with the largest 60 percent too large.
Rausche damping was applied to the same sand example and the results are shown in Figure 10. Forces of 6400 kN, 5960 kN, and 5060 kN were applied. The results for the two larger forces, given by the capacity at zero velocity, are reasonably close to the static capacity. However, the case of the smallest applied force gives a predicted static capacity of about 4800 kN, about 20 percent larger than the given capacity.
This study showed that it was possible to induce Statnamic-like forces using a wave equation analysis by GRLWEAPTMwithout modification for this particular application. The force time record is of the appropriate length and a similar shape. The forcedisplacement curve has the appropriate shape and appearance.
Examples of the modeling of Statnamic tests in clay for the same pile that was analyzed above are given in Figures 11 and 12. Figure 11 shows the results for three Statnamic tests of different applied force levels
8
It is well-known that the Smith damping representation does not produce results that agree with experience. (Rausche et a1 1994) has shown that the Smith damping constant does not agree with measurements for the assumptions used. Some results reported for Statnamic (James et a1 1994) show a character of result that agree with the wave equation analysis. In particular, the case presented with Rausche damping for sand has the observed result. As the applied Statnamic force is increased the agreement between the evaluation method used and the actual capacity is improved. The conclusion has been reached that a substantial permanent set should be achieved.
Fischer, H. C. 1984. Stress wave theory for pile driving application. lecture at the Second International Conference 011 the Application of Stress Wave Theory on Piles, Stockholm, Sweden. Goble, G. & F. Rausche 1970. Pile load test by impact driving. Highway Research Record, No. 333. Goble, G. G. & F. Rausche, 1976 Wave equation analysis of pile driving - WEAP program. Volumes 1 through 4, FHWA #IP-76-14.1 Through #IP-76-14.4. GRLWEAPTM- Wave equation analysis of pile driving 1998. Goble Rausche Likins and Associates, Inc., Cleveland, Ohio. James, M., A. Sy, & R. Campanella 1994. A comparison of statnamic and static load tests on steel pipe piles in the Fraser Delta. Proceeding of the Annual Svniposium on Deep Foundations, Vancouver Geotechnical Society, Vancouver, B.C., Canada. Heerema, E. P. 1979. Relationships between wall friction displacement velocity and horizontal stress in clay and in sand for piled driveability analysis. Ground Engineering. Rausche, F., G. Likins, and G. Goble 1994. A rational and usable wave equation soil model based on field test correlation. Proceeding International Conference on Design and Construction of Deep Foundations, Orlando, Florida, December. Rausche, F., G. Thendean, H. Abou-matar, and G. Goble 1994. Investigation of dynamic and static pile behavior from modified standard penetration tests, Federal Highway Administration Report. Samson, C. H., T. J. Hirschjr., & L. L. Lowery, 1963. Computer study for dynamic behavior of piling. Journal ofthe Structural Division, ASCE, Volume 89, No. ST4. Smith, E. A. L. 1950. Fundamentals of electronic calculation pile driving impact. International Business Machines. Smith E. A. L. 1957. What happens when liammer hits pile. Engineering News-Record 159, September, 5. Smith, E. A. L. 1960. Pile driving analysis with the wave equation. Journal of the Soil Mechanics and Foundation Engineering Divisioii, ASCE, No. 86, August.
The results presented here support the idea that the Rausche damping constant should be studied hrther and that recommendations should be developed for the required damping constants. CONCLUSIONS Curves are presented to assist in obtaining cushion stiffness requirements for concrete piles to limit the induced compression stress at the top of the pile. Impact stresses can be determined for a selected set of ram mass, cushion stiffness, and pile impedance. These values have been found for reasonable ranges of the variables using a closed form solution of the one dimensional wave equation. They have not been checked against field measurements and such checks should be performed. It should also be noted that the analysis does not include the pile cushion and the helmet.
A wave equation study is reported on the modeling of the Statnamic test. The model matches the variation of the induced force at the pile top quite well. Smith and Rausche damping representations have been studied. The Rausche damping representation seems to match the field observed results. In view of the observed poor comparison between the Smith representation and field observations it would be desirable to determine appropriate Rausche damping constants by additional studies.
REFERENCES Clough & Penzien 1975. Dyiianiics of Structures. McGrawHill, Inc., New York. Coyle, H. M. & G. C. Gibson 1970. Empirical damping constants for sands and clays. Journal of Soil Mechanics a i d Foundations Division, ASCE.
9
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Applicationof Stress-Wave Theory to Piles, Niyama & Beim (eds)02000 Balkema, Rotterdam, ISBN 90 5809 150 3
Analysis of bearing capacity of rock-socketed piles based on wave equation theory Lai-bing Cai Fujian Academy of Building Research, Fuzhou, People’s Republic of China
ABSTRACT: Based on wave equation theory, a dynamic loading model is used to analyze the behaviors of rock-socketed piles in this paper. The results of dynamic loading tests are analyzed by CAPWAP. From the analyzed results of twelve dynamic loading tests on rock-socketed piles, the behaviors of shaft resistance of the socket and toe resistance are discussed, a relationship of shaft resistance of socket to the length of socket and the rock strength and the construction of piles is presented. And the parameters of dynamic loading tests are estimated. Some suggestions for dynamic loading tests of rock-socketed are made to analyze the bearing capacity.
1 INTRODUCTION
vanced in this paper. Through analyzed the results of CAPWAP, the bearing behaviors of rocksocketed pile are discussed and a method is given to determine the bearing capacity.
Smith (1 960) initially made numerical analysis on one-dimensional wave equation of pile foundation with the finite difference scheme. From then on, many scholars put up with their numerical models and gave the calculating formulations all over the world, so wave equation analyzing method was widely applied to engineering practice. Basing on the previous research results, Goble et al. (1980) modified the model with better formulation, and put out the program of CAPWAP, which can better simulate the fact of pile-soil interaction. Thereby, it was virtual stage for wave equation taken into engineering practice. At present, wave equation analyses are used to simulate the process of driving pile, judge the feasibility of driving pile and the damage of pile, and determine the bearing capacity of pile. Rock-socketed pile foundations can provide an effective and economical means of transmitting large concentrated structural loads through overburden soils to underlying rock, especially under the condition of overburden soil being soft and rock embedding shallowly. The bearing behaviors of rock-socketed pile change with rock geological condition, construction technology, sediments of pile bottom, rock-socketed depth and so on. It would be taken much manpower and many financial or material resources to determine bearing capacity of rock-socketed pile by static loading tests. Based on the theory of wave equation, a dynamic loading model of rock-socketed pile is ad-
2 DYNAMIC LOADING MODEL During wave equation analyzing, the total driving resistance R can be broken up into two distinct portions: the static resistance R, and dynamic resistance R, R = R , +R, (2.1) The static resistance and the dynamic resistance are represented by +U, u 200
75 -90
Good
I00 - 200
50 - 75
Regular
50 - 100
25 -50
Weak
25 - 50
< 25
Very Weak
1-10
Chart 4 - Toe resistance.
According to Canadian Foundations Engineering Manual (1985) and Ladanyi and Roy (197 1):
4,,
=o,.k,d
(3.3)
: 0.13
0.17 0.20 0.13
0.17
0.20
CT~;M~~
164.0 42.21 56.20 64.94 56,140 73,420 86.370
o,M,,
52.5
13.51 17.67 20.79 17,970 23,500 27,650
35.0
9.01
101.0 26.00 34.00 40.00 34,580 45,220 53.200
in that: Q M ~= maximum toe bearing pressure; X
average unconfined compressive strength cr, = of rock core; K ~ , , = empirical factor; and d' = depth factor.
11.78 13.86 11,980 15,670 18,430
It was considered, in the design phase, that the ultimate capacities of skin fiiction and toe resistance would be the following:
Chart 3 - Skin friction.
3270kN 2 to oc = lMPa, 7Max = 0.25MPa 11,980kN s to oc = 35MPa, qMax = 9.01MPa and ksp= 0.13 (Q,Jp 3270kN + 11,980kN = 15,250kN QUi
Qp =
TMax
(ma)
0.25
0.20 0.25 .0.30
0.20
10
0.63
0.95
8230 10,320 12,420
8
0.57
0.71
0.85
7450
9280
11,100
6
0.49
0.61
0.73
6400
7970
9540
0.79
0.30
5
0.45
0.56
0.67
5880
7320
8760
3
0.35
0.43
0.52
4570
5620
6800
2
0.28
0.35
0.42
3660
4570
5490
1
0.20
0.25
0.30
2610
3270
3920
1 8
--I
k,
1 5
I -
It has been verified that the safety factor for a work load of 4400kN is 3,4, and therefore satisfactory. 4 HIGH DYNAMIC LOAD TEST 4.I Pile and Instrumentation Features
The conditions of construction and the test, are represented in Figure 3.
(3.4)
443
Figure 3 - High Dynamic Load Test - Typical setup. The pile was cast in place with the concrete fck=25MPa (characteristic compressive strength) and heavy longitudinal steel reinforcement associated with spiral reinforcement cage. Instrumentation for the Dynamic Load Test (DLT) and the test itself followed the recommendations of ASTM D-4945/89, AASHTO designation T-298/93 and ABNT-NBR 13208/94 (Brazilian Standard). Constructively, the shaft space incorporated between the confining soil and the top of the alluvium of basalt is constituted by a steel casing with a 3/8" thickness and an internal diameter of 1,20m driven with a casing oscillator rig. In order to carry out the test, a supplementary 47cm shaft of concrete was executed above the steel casing, where the conditions of longitudinal reinforcement were maintained according to project, altering the spacing of the spirals to 8,Ocm. The excavation of the rock was accomplished with the use of explosives and a compressed air system, for the purpose of expulsing underground water which originates in the hydraulic load of the Uruguay River. After the excavation in rock, the pile was concreted to a quota of 333.62, corresponding to Figure 3.
The instrumentation proceeded through the installation of four strain transducers and four accelerometers, positioned every 90 degrees, in a manner to minimize the possible effects of bending resulting from the dynamic impact. In the case of the accelerometers, two piezoeletrics and two piezoresistives were used. The opening of four "windows" became necessary in the steel casing, for the installation of the sensors in the concrete nucleus, as represented in Figure 4. Chart 5, which follows, specifies the characteristics of the tests. Chart 5 - Test features. COLUMN = P5 BORED PILE = T6 SOCKET
4
LENGTH
MU.
ROCK
120
130
(tD
3.20
440
_. .
444
TOTAL
(4
*Maximum worklng J
LENGTHS (m)
16.29
BELOW
BELOW
SENSORS
GL
15.01
14.01
4.2 Hammer Description
- The hammer weight should, at least, be equal to
1.5% of the anticipated static test mobilized load;
Dynamic Load Test on cast in place piles with high load capacity have had routine application in the USA, Europe, Asia etc. and in Brazil. The technical literature available gives an account of several cases of dynamic tests on cast in place piles, as discussed by Rausche and Seidel (1984), Balthaus et al. (1985)Townsend et al. (1991), Seidel et al. (1996), Sawai et al. (1996) etc. The option of using the high strain dynamic load test (DLT) instead of the static load test, which is routine in many places, comes from the attractiveness of the economics and the scheduling benefits. The cost of the high strain dynamic test in Brazil is 25% less than the cost of a static load test (SML-slow maintained load), and represents a reduction of 85% in time consumption. The device designed to apply high impact forces on the head of the bored pile consists of the free fall self-propulsion hammer with special features to test piles with high load capacity and variable diameters, varying from 70cm to 250cm. The selection of the proper hammer size is essential for successful high-strain dynamic testing hammer weight, and drop height. The cushion details must be appropriately chosen so that hammer impact causes sufficient pile movement in order to mobilize the required soil resistance, and to assure that dynamic stresses in the shaft will not impair its structural integrity. The hammer apparatus must be constructed in a way to facilitate mobility around job sites and assure a uniform impact to the head of the pile (Hussein et al., 1996). The free fall hammer conceived and in use in Brazil consists of the following main components, shown in Figure 3 : the steel striker plate, the cylindrical - 2 pieces - steel cage (both of these components being referred to as the hammer), the set of round 20kN steel rams allowing modulation up to 200kN in maximum drop of 3rr4 and the hydraulic system to lift the ram by an automatic catcher, which activates and controls the hydraulic jacking equipment. The steel striker plate consists of a round 2.1m diameter by 100mm thick steel plate, with four lateral latches that adjust themselves according to necessities in relation to the diameter of the pile to be tested, and functions as a hammer support on the pile top and as a helmet.
- The hammer drop height should be approximately
8.5% of the pile length, with a minimum value of 2.0m; and - A plywood cushion thickness of t = L2 / 2 0 , where thickness ( t ) is expressed in and pile length ( L ) and pile diameter ( D ) are in meters, a minimum value of lOOmm is suggested, with an additional 15Omm being added when the pile length exceeds 30m. In this particular case, it was postulated that a ram of 120kN, with a maximum free fall height of 2 . 5 ~ and a plywood cushion thickness of 150mm by 1.l m diameter would be used to comply with the recommendations. The diameter of the plywood cushion should be about 80 to, at most, 90 percent of the shaft diameter in order to be able to centralize the blow. The thickness of the steel striker plate to be placed above the top of the pile should range from 50mm to 100its minimum diameter being equal to the diameter of the cushion the ram, ideally, is a simple drop weight which can be raised to a variable height. In order to facilitate, rams can be made in segments which can be assembled at the job site. The ram to be used needs to be “guided” to assure that the pile top and ram bottom are perfectly parallel during impact, in order to assure uniform contact stresses during impact. Between the steel striker plate and the top of the pile a 20mm thick steel wired rubber pad was installed with a 1.0m diameter in order to evenly distribute compressive stresses on the pile head, absorbing the dynamic impact, and preserving the structural integrity of the pile top above the electronic sensors, which measures data to be associated with the test.
4.3 Technical Criteria
To determine the hammer weight, drop height, and plywood cushion thickness to be used in this p&icular test, the bibliographic recommendations of Likins (1994) and Hussein et al. (1996) were followed:
Figure 4 - Details of the windows in the steel casing for the installation of the accelerometers and strain transducers
445
4.4 Case Method and thefield DLT
The test was done by a sequence of blows with variable drop heights as a dynamic load test with increasing energy. References can be seen by Aoki (1989), Niyama and Aoki (1991), Hussein et al. (1992), Aoki and Mello (1992), Mello and Aoki (1993), Beim and Aoki (1996), Mello and Paraiso (1998), among others. It’s important to note that the sequence of blows was done specifically on a concrete shaft inside in steel case. The DLT with increasing energy can be seen as a cyclic test similar to a static load test with increasing load steps where each loading cycle corresponds to an increasing energy impact. The test started with a drop height of 0.5m up to 2.5m with increments of 0.5m. The input data supplied to the PDA-PAK95 equipment for the beginning of the DLT were: e SP (pile specific weight): 24kN/m3 e AR (pile area on the sensors level): 11,309.8cm2 EM (pile dynamic modulus): 3 12tf7cm2 e EA/C (pile impedance): 1,007.31tfXm e WS (wave velocity): 350Ods 0 JC (Case Damping factor): 0.5 (medium value) At the same time precise millimetrical set measurements were collected from the shaft of the pile for each blow. Figure 5 presents the force and velocity x EA/C traces as function of time for each blow, obtained from the strain transducers and accelerometers processed by Case Method and reprocessed by Datpro Program. These traces exhibit good proportionality between force and velocity, and the peaks of force show an increase as the energy from the dynamic impact rises, which is perfectly consistent.
Figure 5 - Force and Velocity traces. developed internally in the steel casing and between the steel casing and clay soil, as following: - drop height; H FMX = the maximum compression force; EMX = the maximum energy transmitted past the transducer; %EMX = efficiency; SET = permanent SET; DMX = the maximum downward displacement at the transducer location; total mobilized resistance; toe resistance; lateral skin friction; lateral quake; toe quake; Damping factor; lateral damping factor; and toe damping factor. 5 CONCLUSIONS
4.5 Cupwup Analysis In the Capwap analysis the pile was initially modeled considering it’s uniform cross section. However, in the case of the pile tested, it was necessary to model impedance variations along the depth into the rock for improving the “match” between the curves of force and velocity measured and calculated. In this particular case, the analytic procedure was extremely necessary. Chart 6 summarizes the results of the Capwap analyses for each blow applied on the top of the pile. Chart 7 shows the typical results from the Capwap Analysis with reference to all five blows. It considers the values of skin friction, in the shaft space, inserted in the alluvium of the basalt,. between quotas 321.00 and 317.80, and toe resistance disregarding the portion of skin friction 446
Having the results presented by Capwap Analysis and Case Method as a reference it is concluded that: 0 The traces of force and velocity x EA/C presented in Figure 6, shows with clarity the increase of the load mobilized with the applied growing energy. 0 The liquid energy transferred to the pile grew with each applied blow, arriving at 29% in the fifth blow, a low value for the hammer in fi-ee fall without the interference of mechanical friction. The reason of the low transfer of energy is in consequence of the great loss of energy with the shock absorber system used. e It is verified that the maximum mobilized resistance of friction and point are manifested respectively when the fourth and fifth blows and confrmed with the obtained values of Qs and Qt according to Chart 6.
Chart 6 - Capwap analysis - general data.
-
BLOW
Rs SKIN up to 321.00
(kN)
Level (kN)
Rs-SKIN between level 321,OO and 3 17.80 (kN)
1840
1661.0
179.0
KS
'I (m)
R p Point Resistance (kN)
4025.0
RU Total ReRU sistance RU (Rs + Rp) (kN) (kN)
FMX (kN)
EMX (kN.m)
4030.0
9.0
% SET DMX EMX (mm) (mm)
~
1
0.50
4204.0
5865.0
15.0
0
2.72
2
1.00
4572
3148.5
1423.5
5101.0
6524.5
9673.0
5750.0
20.0
16.7
0
3.80
3
1.50
6621
361 1.5
3009.5
5180.0
8189.5
11,801.0
9020.0
37.0
20.6
0
4.85
4
2.00
6365
2938.0
3427.0
7050.0
10,477.0
13,415.0 12,460.0
61.0
25.4
0
6.21
5
2.50
4114
1387.0
2727.0
11.129.0
13,856.0
15,243.0 16,420.0
86.0
29.0
0
7.45
Chart 7 - Capwap analysis - socketed in rock. H (m)
RU (kN)
RP (kN)
1
0.50
4204.0
4025.0
2
1.00
6524.5
5101.0
BLOW
Ks
SMITH
Qs (mm)
(mm)
179.0
1.377
1423.5
2.280
(kN)
Qt
J CASE CAPWAP
Js (s/m)
J, ( s h )
1.030
0.640
1.754
0.87
1.750
1.ooo
1.165
0.25
3
1.50
8189.5
5180.0
3009.5
3.002
2.587
0.701
0.876
0.44
4
2.00
10,477.0
7050.0
3427.0
4.158
3.900
0.7 13
0.75 1
0.52
5
2.50
13,856.0
11,129.0
2727.0
4.000
4.682
1.136
0.426
0.41
It is observed that the measured value of zero set in all the blows leads to the consideration that the pile still presented an additional point load to be mobilized, however, the value of the load mobilized in the fifth blow indicated factor of safety larger than 3, in relation to the predicted working load, being opted by the paralyzation of the test. 8 It is considered that the ultimate load mobilized in the dynamic test can be interpreted in the following way: - Ultimate Bearing Capacity (5th blow) = 13,856kN - Ultimate Bearing Capacity, considering the friction resistance (4th blow) added to the point resistance th blow) = 14,556kN The test showed the saturation of the skin friction in the 4th blow and maximum point mobilized in the 5th blow, whose obtained values present a close approximation with those considered in the semirnpiric analyses in the project phase. The dynamic test in piles of high load capacity has become a reliable tool and has reduced costs and shortened evaluation deadlines for the quality of this type of deep foundation, as long as the hammer to be used presents the available potential energy to the load to be mobilized.
6 ACKNOWLEDGEMENTS The authors wish to thank the support given by the Company CBPO-Grupo Odebrecht,and Mello de
Azevedo contractor responsible for the construction of the bridge on the Uruguay River and, especially, to engineer Claudio Menin, project planner and responsible for the structural calculations of the bridge.
7 REFERENCES AASHTO T 298/93 - American Association of State Highway Officials Standard - Method of test for high-strain dynamic testing of piles. ABNT NBR 13208/94 - Associaggo Brasileira de Normas TCcnicas - Estacas Ensaio de Carregamento Dinilmico, in Portuguese. ASTM D4945/89 - American Society for Testing and Materials Standard - Test method for highstrain dynamic testing of piles. Antunes Sobrinho, J. et al. - 1999 - Criterios de projeto na construqgo da barragem da UHE Ita. XXIII Seminario Nacional de Grandes Barragens. In Portuguese, Belo Horizonte, Brazil. Aoki, N. - 1989 - A new dynamic load test concept. Proc. Discussion Session on Pile Drivability, XI1 ICSMFE, Rio de Janeiro. Aoki, N. & Mello, V. F. B. - 1992 - Dynamic loading test curves. Proc. Fourth International Conference on the Application of Stress-Wave Theory to Piles. Hauge, Holland. Balthaus, H. G. et al. - 1985 - Dynamic load test 447
German practice. Proc. 1lth Int. Cod. Soil Mechanics and Foundation Engineering, San Francisco, USA. Barton N., Lien, R. & Lunde, J. - 1974 - Engineering classification of rock masses for the design of tunnel support. Rock Mechanics, 6 (4), 189-236. Beim, J. W. & Aoki, N. - 1996 - Dynamic load test method with variable energy. Proc. Fifth International Conference on the Application of Stress-Wave Theory to Piles, Orlando, USA. Beim, J. W. and Paraiso, S. C.. - 1992 - Dynamic testing of enlarged base Franki piles. Proc. Fourth International Conference on the Application of Stress-Wave Theory to Piles, Hauge Netherlands. Bieniawski, Z. T. - 1984 - Rock mechanics design in mining and tunnelling. Balkema, Rotterdam. Broms, B. B. et al. - 1988 - Bored in residual soil and weathered rocks in Singapore - BAP I Ghent - Belgium. Canadian Foundation Engineering Manual - 1985. Capwap Users Manual - 1996 - Goble Rausche Likins and Associates, Inc. - Cleveland, Ohio, USA. de Mello, L. G. & Paraiso, S. C. - 1998 - Variable energy dynamic load test on a 1.0m diameter CFA pile. Proc. Belgium - BAP 111. de Mello, V. F. B. & Aoki, N. - 1993 - Updating realism on large diameter bored piles. Deep Foundations on Bored and Auger Piles, Ghent, Belgium. Hussein, M. et al. - 1992 - Dynamics of pile driving as a function of ram drop height. Proc. Fourth International Conference on the Application of Stress-Wave Theory to Piles. Hauge, Holland. Hussein, M. et al. - 1996 - Selection of a hammer for high strain dynamic testing of cast in place shafts. Proc. 5'" Int. Cod. on the Application of Stress Wave Theory to Piles, Florida, USA. Infanti, N. et al. - 1999 - Tensees residuais nas obras subtenheas da UHE Ita. XXIII Seminario Nacional de Grandes Barragens, Belo Horizonte, Brazil. Lima, M. J. C. P. A. - 1983 - Prospecqiio Geotecnica do Subsolo - Livros Tecnicos e Cientificos Editora S. A., Rio de Janeiro, Brazil. Niyama, S. & Aoki, N. - 1991 - Correlaqilo entre provas de carga diniimica e estatica no campo experimental da EPUSP/ABEF. In Portuguese, 2rd -Seminario de Engenharia de FundaGdes Especiais e Geotecnia, SEFE, Siio Paulo, Brazil. Rausche, F. & Seidel, J. P. - 1984 - Design and performance of dynamic tests of large diameter drilled shafts. 2"d Int. Conf. on the Application of Stress Wave Theory on Piles, Stockholm. Rausche, F. et al. - 1985 - Dynamic determination of pile capacity. Journal of Geotechnical Engineering, Vol. 111 - Nc 3 - ASCE: 367 - 383.
448
Rausche, F.- 1988 - High strain testing of drilled shafts. Seminar for dynamic testing of cast in situ piles for capacity and integrity - Boulder, CO. Seidel, J. P. et al. - 1996 - Dynamic testing of barretes for a cement silo project. Proc. Fifth International Conference on the Application of Stress-Wave Theory to Piles. Orlando, USA. Townsend, F. et al. - 1991 - Dynamic load testing of drilled shafts - Final Report, Univ. Florida, Gainsville, USA. Zhang, L. & Einstein, H. H. - 1998 - End bearing capacity of drilled shafts in rock. journal of geotechnical and geoenvironmental engineering, vol. 124.
7 SPT measurements and special field monitoring test
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Application of StreSS-WaVe Theory to Piles, Niyama & Beim {eds)02000 Balkema, Rotterdam, ISBN 90 5809 150 3
Keynote lecture: Frequency characteristics of stress wave and penetration during SPT K. Fujita Department of Civil Engineering, Science University of Tokyo,Noda, Japan
ABSTRACT Recently wave equation analysis has been successfully applied to the SPT to determine the energy transfer ratio for adjusting the N-value, but should not limited to this purpose The SPT N-value has been applied to the design of foundations for more than SO years, however, the correlation between the Nvalue and the design parameters have been estimated crudely based on experience, so that new methods or constitutive equations should be established for determining the design parameters rationally. The frequency characteristics of the stress wave and the behaviour of the penetration of the rod system are introduced, because the relation between them could be keys to hrther studies 1 INTRODUCTION
As the ISSMFE was unable to standardize the method and equipment to be used for the SPT internationally, the SPT Working Party of the ISSMFE Technical Committee on Penetration Testing proposed the “International Reference Test Procedures on SPT”, at ISOPT-I, Orland in 1988. A method using stress wave equation analysis was introduced in Document 4- Calibration Methods for the adjustment of the N-values in the SPT to account for different magnitudes of driving energy, as a means of quality control for the SPT. The method was established by referring to studies made by Uto et a1 (1974), Schmertman et al. (1979) and others In the author’s opinion, the calibration methods are unable to account for a difference in the size of equipment, however, this could be corrected by calculations and the formula to be applied for the adjustment of N-values should be revised. Matsumoto et al. (1992) have applied the method of two-point strain measurements proposed by Lundberg & Henchoz (1977) to the measurement and analysis of a SPT experiment carried out in the field The displacement-time diagram, the transmitted energy-time diagram and stress-time diagram were verified, and the calculated data corresponded well with the measured data Dynamic measurements of the energy transfer ratio have been carried out in practice in USA and other countries with increasing frequency. Aboumatar et al. ( 1 996) have presented a typical dynamic SPT records including the measured force, adjusted velocity and displacement with time computed by
integrating the velocity, which was equal to the observed final set as shown in Figure 1 The wave equation soil constants, i e the quake, damping constant and mass for use i n the wave equation analysis, as well as the static soil resistance obtained from the calculated toe force and motion were determined reasonably well Abou-matar & Goble ( 1997) have contributed a valuable paper which broadly discusses the problems associated with the dynamic measurements and analyses concerned with the SPT They pointed out that changes in the cross section of the drill rod may cause substantial changes in the measured Nvalue and presented method for detecting and avoiding various problems based on onedimensional wave mechanics Fujita & Ohno (2000) introduced studies concerning the application of the stress wave theory to the SPT in Japan, and included a figure that indicates that the SPT hammer strikes the rod more than three times afier the first impact They suggested that analysis of the frequency characteristics of the measured stress wave would provide valuable information for the study of the SPT concerned The term = f i known as “wave propagation speed’, etc is applied in one-dimensional wave mechanics and assumes that the rod length is infinite or the rod area (diameter) is negligibly small The wave propagation speed in the anvil, rod connector or hammer is about 60 to 70% of c So, the analysis should be take into consideration the frequency of the wave
451
h
E E 4 c
5 E
y
-cac 6 d 2
-44-
-
\'elocity.(LA,c) (IrN) --Divplaceinent
P O I C(hN) ~
0 CL
I
I
I
I r)
I
I
-4
(mtn) I
I
40 Time (ms) Figure 1 Dl ii;uiuc SPT ineasureinent records - force, L elocih and displacement (aficr Abou-matar et a1 1096) 0
\
70
30
II
Final penetration
netrdtiori vs time l'enetratiori
.Maximum penetration
I 10
20
30
40
50
GO
'50 I0
Time ( ms)
Figurc 2. Strcss and pcuctration L'S. tiinc diagram. (Fujita 2000).
In this paper, the penetration or vertical movement of the rod, the frequency characteristics of the stress wave and the relation between the rod penetration and the frequency of the stress wave will be introduced, while the study is still proceeding. 2 PENETRATION VS. TIME DIAGRAMS Figure 2 is an example o f a diagram showing records of the penetration of the rod with time and the stress in the rod with time (stress wave) measured near the top of drill rod during one blow of the SPT. The figure shows that the hammer strikes the rod again (second impact) about 14ms after the hammer blow (first impact) and the penetration increases accordingly. The penetration at the time of the second impact is smaller than the maximum values 452
of the penetration which occurs between the first and second impacts due to the rebound of the rod and/or ground before the second impact. In this case, during the completion of the SPT there are at least two impacts of the hammer, and this evidence should be considered in the wave equation analysis. Figure 3 is a diagram summarizing the measured penetrations of the rod as a finction of time during the part of the SPTs at a depth of l l m . The maximum penetration and the final penetration for each blow are around 12 to 13mm and 4 to 7mm, respectively, and the rebounds of the rod and/or ground are between 6mm and 8mm. While the final penetrations (set) per blow which contribute to the N-value are scattered, the maximum penetration per blow is almost identical in all cases. Attention should be paid to this from the stress wave analysis point of view, because the energy
Figure 3 Penetration 1 s time diagram - at a depth of 1 l n i (Fujita 2000)
case shown in Figures 3-4, in ground which consists mostly of sand and gravel layers of alluvial and diluvium strata, and where the water table is 1 10m below the ground surface and the N-value is smaller than 50 Figure 4 shows the relation between the maximum penetration and the fhal penetration of the rod, (a) on a linear scale and (b) on a logarithmic scale measured at various depths during the SPT, without adjustment for the energy transfer ratio The energy transfer ratio varies from blow to blow in this case, mainly between 0 6 and 0 9, with an average of 74 Values of the maximum penetration smaller than 12mm were not observed during the SPT at this site, except for the case when the energy transfer ratio is small When the maximum penetration is about I 3 to 14mm,the final penetration will be less than about 1 Omni When the maximum penetration is larger than about 22mn1, the difference between the maximum penetration and the final penetration is less than about 21nm The maximum penetration occurs after the time (2L/c), in this case 3 FREQUENCY CHARACTERISTICS OF
STRESS WAVES The straidstress with time measured by strain gauges installed near the top of the drill rod as a function of time is called the “stress wave” in this paper, and an example was shown previously in Figure 2 In the wave equation analysis of the SPT, the hammer, guide rod, anvil, drill rod, coupler and sampler are generally assumed to be made of the same material and to have the same cross-section. except for the hammer The downward travelling wave arid upward travelling wave caused bv the hammer blow are propagate in the rod in the downward and upward directions, respectively, at the wave propagation
Figure 4 Relation betncen the iiia\iiiiuiii pcnetratioii and final penclration (Fiijita 2000)
transmitted to the rod is likely to correspond to the masinium penetration, not only to the final penetration So, dynamic measurements of the penetration (vertical displacement) should be made The strain or stress and the penetration of the rod were measured about 2 to 2 5m aboveground, near the top of the drill rod and below the anvil, in the 453
= . i i 111 ~ for the material, based on speed c- = ,/A one-dimensional wave dynamics However, the speed depends on the shape of the members, and decreases when two-dimensional wave dynamics are applied For example, the speed in the hammer or anvil would be about 70% of the one-dimensional wave propagation speed for the material The waves are partially reflected at the ends of the rod system and at the junctions between the different cross sections so that various waves are propagated in the rod following the hammer blow In order to examine the frequency characteristics of the stress waves, the method of Wavelet transformation was chosen, with which it is possible to obtain a thi-ee-dimensioiial diagram giving the frequency, time and power spectra Figure 5 shows two frequency vs power diagrams obtained by Wavelet transformation of stress waves, measured at (a) a depth of3m and (b) a depth of 17m The natural frequencies of the rod systems employed at (a) and (b) are 0 286kHz and J>
t
i I
I
I
0 1 ISkHz, respectively, assuming the wave propagation speed is = Jty. while the predominant frequencies shown in Figure 5 (a), (b) are calculated to be 0 19kHz and 0 074kHz, respectively, which are about 65% of the natural frequencies In Figure 5 , more than four prominent frequencies can be seen including the predominant frequency, these peaks are referenced to as A, B, C, D ..., from low to high frequency The natural frequency is located between the frequencies of the peaks B and C According to the time-power diagram (not included in this paper), the peak B occurs about the time of the second impact between the hammer and the rod As shown in Figure 6, the reciprocal of the time between the first and the second impacts corresponds well with the fiequency at the peak B 4 RELATION BETWEEN POWER AND PENET RATION Figure 7 indicates that the power at the peak A corresponds to the final penetration of a hammer blow of the SPT The frequency at the peak A is estimated to be around 15 to 20Hz, which is probably determined by the duration of the rod movement of 50 to 70 ms, as shown in Figure 3 The unit of power is ((N/mm2)s)2,where the .s is included in the term is the data sampling inteival, 5 s in this case The power is an index provided i n the Wavelet transformation, which is probably related to the penetration and the stress in the rod as well as the dynamic and static soil parameters, however, it has not yet been clearlv identified in the study ofthe SPT
I
Pztietration 71nm
001
01 1 Frecluel1cy (hHz)
10
(b) at a depth of 171n figiirc 5 Pciictration 1 s poucr spcclrum (Fqita 2000)
Figure 6 Reciprocal of the tiinc bctnceii 1 “ and 2’ld 1111pacts plottcd agaiiist fiequcnc! at peak B (Fujita 2000)
454
nicnsiircnieiils on SPT. llriic. .>" Iiii. 'oil/. n i l , l ~ ~ p l ~ c o i i o ~ oj'.sfrf~.~.~-lCnl~f~ ?%lV)i?' if, I'ilC.V, Orlniltl. 16.3- 175.
Figtire 7. Penetration plotted against power st pcak A. (Fujita 2000).
Abou-inarar. H.. X: G. Goblc. 1997. SPT dynamic analysis and tnc:isiircincnIs. .1. o/( iGK . I Fuiita. K. 2001). Strcss vavc i k o n iipplicilion 10 standard pcnctr;itioii ICSI in japan. /+iic. c;" iiii. ( ' o i i t iiiJ .i/~p/icii/ioi7 ~/.S/rc.v.v-~rmv /h<wcl'/oPilc,s, S m fmilo. (to be publislicd) ISSMFE Tccllnical Coininittcc on Pcnctratioii Testing-SPT \\.orking party. 1988. Standard Pcnclration Tcst (SPT). Irircrt~:itionnl rcfcrciicc tcst proccdurc. Iri Dc Kuilcr (cd). /'c~I1~~ir(J/IOIl 7?s/iilg I9HX. 0 r I ~ i I d: ( 1 ).7-26. Rotterdam : B;i Ikcnia. Lundbcrg. B.. X: A. Hcnchoz. 1977. Analysis o f el;islic w i ~ c s froin two-point slmin mcasurcmcnl. .I. of' I*~.qurii~rei~/. .\/N,. Nd> N , N(J>N, Y anase (1963) N(s>NLV N ( J > N , Nct>N,br Yaniacla et al. (1004) Ni>N,, N(s=N,. NisN,, Ncs'N, NcrNational Research Council, Washington, DC. Vol. 1 pp. 1-51. ISBN 0-309-04613-0 ONeilI Michael W., Vipulanandan, C., Wong, D., (1990). "Laboratory modelling of vibro-driven piles". Journal of Geotechnical Engineering, ASCE, Vol. 116, No 8: pp. 1190-1209 Roberlson, P.K. and Wride, C.E. (1998). "Evaluating cyclic liquefaction potential using the cone penetration test", Canadian Geotechnical Journal, No. 35, pp. 442-459. Rodger, A.A. and Littlejohn, G.S. (1980). "A study of vibratory h v i n g in granular soils". Geotechnique, Vol. 30, no 3, pp. 269-293. Seed, H.B. and Idnss, I.M. (1970). "Soil Moduli and Damping Factors for Dynamic Response Analyses." Earthquake Engineering Research Center, College of Engineering, University of California, Berkeley, Report No. EERC 70-10. Seed, H.B. and De Alba, P. (1986) YJse of SPT and CPT Tests for Evaluating the Liquefaction Resistance of Sands" Proc. INSITU 86, V A 22 p. Schmid, W.E., (1969). "Driving resistance and bearing capacity of vibrodriven model piles". American Society of Testing and A4aterials Special Techn., Publ. 444, pp. 362375 Schmid, W.E. & Hill, H.T. (1967). "A rational dynamic equation for vibro driven piles in sand". Symp. Dynamic Properties of Earth Materials. New Mexico Universityy:p. 349. Smith E.A.L., (1960). "Pile-dnving analysis by the wave equation". Jourrzal of the Soil A4echanics and Foundatioris Divisions, ASCE, Vol. 86, August 1960. Vilung, K., (1997). "Vibratory dnven pile and sheet piles - a literature survey", Report 3035, Div. of Soil and Rock Mechanics, Rqyal Institute of Technology, Sweden, 75 p. Vucetic, M. and Dobry, R. (1988). "Degradation of Marine Clays Under Cyclic Loading. ASCE Journal of Geotechnical Engineering,Vol. 114, No.2,pp. 133-149. Vucetic, M. and Dobry, R. (1991). "Effect of Soil Plasticity on Cyclic Response." ASCE Journal of Geotechnical Eiigiiieering, Vol. 117, No. 1, pp. 89- 107. Vucetic. M. (1993). "Cyclic Threshold Shear Strains of Sands and Clays", Research Report, UCLA Dept. of Civil Engineering, May 1993. Vucetic, M. (1994). "Cyclic Threshold Shear Strains of Sands and Clays", Paper in print, ASCE Journal of Geotechnical Engineering Wamngton, Don. C.: (1989). "Driveability of Piles by vibration". Deep Foundations Institute 141h Annual members Con$, Baltimore, Maryland, USA pp. 139-154. Westerberg. E.. Massarch, K. Rainer. Eriksson, K., (1995). "Soil resistance during vibratory pile dnving". Proc. to hit. Symposium on Cone Penetration Testing, Linkoping, Sweden, Vol. 3, Report 3.41, pp. 241-250. You4 L.T. (1972). Compaction of sands by repeated shear straining. Journal of the Soil Mechanics and Foundaiioiw Division, 1972, Proc. ASCE, Vol. 98, SM7, pp. 709-725.
494
Application of Stress-Wave Theory to Piles, Niyama & Beim (eds) (c> 2000 Balkema, Rotterdam, ISBN 90 5809 150 3
Computation of ground waves due to piling C. L. Ramshaw, A. R. Selby & P. Bettess University cfDurhanz, U K
ABSTRACT: Pre-formed piles are normally installed into the ground using high energy impact or vibratory hammers. Part of the energy is absorbed in advancing the pile, and much of the remaining energy is transmitted into the ground in the form of outgoing vibrational waves. The current study is towards developing a computational procedure to model the outgoing waves, to calibrate the waves against existing site measurements, and then to expand the model by including simple structural forms so as to estimate structure response, including dynamic soil-structure interaction. This paper reports the stage of the work which generates ground vibrations and some calibration against site data. For impact hammers, the analysis has to be tackled in three phases, of hammedpile impact, pileisoil interface effects, and finally the transient outgoing P, S and R waves. For vibro-driving a different procedure is required. A two stage approach makes an estimate of rigid body oscillation of the pile within a localised soil framework, and then the boundary disturbance is applied to an analysis of the wider soil area for computation of the sinusoidal wave disturbances. Two case studies are presented. must first be calibrated against a substantial database of site records of “green-field” vibrations (Uromeihy, 1990), which is the subject of this paper.
1. INTRODUCTION Pre-formed piles, of steel, concrete or timber, are widely used for transmitting building loads from ground surface down through weak soils to more competent soil or rock strata, while interlocking sheet piles are used for temporary or permanent retaining walls. Pile installation is achieved using high energy vibratory or impact hammers. Some energy is lost into the ground as outgoing P, S or R waves. These vibrations may disturb neighbours (Head & Jardine, 1992), and cause cosmetic damage to nearby structures (Wiss, 1967, Selby, 1991, Todd, 1994). Analysis of risk from vibrations has been attempted by empirical estimates of free ground vibrations and arbitrary vibration thresholds for building types (Attewell & Farmer, 1973, Eurocode 3 part 5 , 1996) with some recognition of frequency effects. Structural analysis for vibration enhancement has been attempted in isolation.
The use of finitehfinite elements to generate ground waves has been demonstrated (Ramshaw et al, 1998), using the package ABAQUSO, and a relatively simple energy transfer at the pile-soil interface. This has now been extended, so as to model the energy input from the hammer, the stress waves in the pile, and the subsequent outgoing ground waves. However, because of the complexity of the system, and the need to use small nodal spacings (xsin(wt + $)
(3.2)
The amplitude of the movement of pile head in the next calculations will be found as bellow:
A ( @ )= W(H;(01
awiax=e
(3.9)
(3.3)
3. I Meckanical characteristics ($the soil and the pile
Sinusoidal case:
?'he pile is formed of fixed sectioned elzments. However the soil is a superposition of the homogenous layers Figure 3. The elements [O,H], situated along of the pile: xo = O l x ,
<x,
T ( H , t ) = Fsinmt]
External loads
By forming of a global transfer matrix the solution of the system is reached through:
(3.13)
S ( H ) = T S(o)
M(H) =0
(3.14)
T ( H )= F
4 EVALUATION OF THE CALCULATIONS AND EXPERIMENT CONCLUSIONS
The coefficient of subgrade reaction value at J, interval.
k I( ” j = kl -p,w’
The experiment has been done on a partially embedded in the soil, having fixed sections composite pile D=l m. The part of the pile, situated in the soil, is J , , and its situated on the soil part is J , .
(3.15)
The equations number (3.10)-(3.14) has to be solved.
The part, situated in the soil: 3.4 Solution oftlie system in the interval [0,H]
I , = 8 . 2 5 m , p I = 3 . 1 x I 0 3 k g n - ’ , r , =5x10’ NI,?’
The vector “S” is separated to two sub-vectors, which provides the limit conditions at the pile end.
The part, situated on the soil:
s’=/
(wl,el,M ’ J ) , S”=’( w 2 , Q 2 , M 2 , - - T 2 ) ,
The “ k , ” value in the calculation has been chosen
The pile end limit conditions; embedded pile -I
w (0) = O,G‘(O) = O,G’(O) = l,-?(O)
-1
w-(o)= O,%’(O)
in a manner, permitting the appearing of the effects of the characteristics of the soil against the pile movements.
= 0, (3.17)
= O,M2(0)= O,-T2(0) = I,
k~
Each solution of the equations number (3.10)(3.14) is a linear combination.
2.5x107~
3x10‘1
5x10’
N117.2
Flexibility - frequency curve:
a’s’ i- a’s2
(3.18) The a’ and a’ constants will be the solution of the system.
= z x l o i j
s = A / F( F = l ) Horizontally displacement-frequency curve:
A = 1000’S 1 4 ~ ’ (3. 8, For the “a”values, which do not make “0” the determinant of this system:
Speed-frequency curve:
’
=10003S/4n’
Acceleration - frequency curve:
r =1 Through the solution of the equations number (3.10) (3.14) the amplitude of movement at the head of pile is calculated as:
-
-
--
A = w(H)= F
-
w2 ( H ) M I ( N )+ W ’(If)%? -,
( H ) (3.20) T - ( H ) M ’ ( H )+ T I ( E . I ) M 2 ( H )
The calculations are made by transfer matrices. At the first stage at each Jl = [ x ~ - ~ , xspace ~ ] the z,elementary transfer matrix is set up,
S. = 2, .Si-l, i = 1,...n is calculated. Then, T = z,?zJI-,
0 0 d / 4 ~ ~
These values are drawn in Figures 4, 5 , 6, 7. The initial slope in the flexibility curve is horizontal, however, this slope in the displacement curve is high. This state of the flexibility curve is demonstrating only that the forces, which are the counterpart of the small frequency values of the used vibrating apparatus may be neglected. It may be said that the speed is constant on the speed curve between 5-10 Hz, and that this fact is depending to the choseii mechanical characteristics. At the change of frequency in value of 10 Hz, the flexibility, displacement, speed and acceleration values are changed in counter way according to the chosen coefficient of soil reaction. This fact shows that how high is the resistance of the soil, so important is the movement of the pile head.
514
FLEXIBILITY LEGEND
v Experitnental values
-
Unstable measunngs Theoretical vzlues
Figure 4. Flexibility curve (k=2500;5000 kdaNm-') Figure 8. Comparison of the theoretical and experimental results.
Figure 5. Displacement curve of the pile head (k=25 00 ;5000 kdaNm-')
I
I
",I1
The flexibility frequency curves, shown at the been drawn using the figure 8 have k=1000,2500,5000 IcdaNnz-' (@2) values. The obtained experimentally flexibility values have been placed on these curves. As it may be seen at the figure 8, the values 2500 kdaNni-2 obtained as a result of the experiment, are shown on the curve. As a consequence the coefficient of soil reaction value of the soil is 2500 ldaNnz-2(tf / m2) As a conclusion, the numerical model and the experiment result fit between 1 Hz and 1,5 I-Iz. The fact that the experiment results between 1,5 Hz and 4,5 Hz are spread is caused due to tlie lack of the resonance at 3,5 Hz. The hai-tnony between the model and the experiment result after 5 Hz obligate to take into account the dynamic effects in the numerical model. Thus it is observed that the iiumerical model which is out of the resonance space is in agreement with the experiment results. 5 CONCLUSION
Figure 6. Speed curve of the pile head (k=2000;2500;3000;5000 kdaNm-')
Figure 7. Acceleration curve of the pile head (k=2000;2500;3000;5000 kdaNm-')
The study is an evaluation of the fitting of tlie coefficient of soil reaction value, obtained experimentally by a vibrating apparatus, mounted at a composite pile head, embedded partially in the soil, with a numerical model. The vibrating apparatus, fixed at the pile head, is applying during the experiment a measurable siiiusoidal horizontal force to the head of pile. The horizontal displacement, which is the equal of this force is measured, too. The horizontal displaccinent flexibility values, which are the equal to the unit force, are easily calculated in place. Using these values, in the calculations, which do not include the dynamic effects, the not-sure coefficient of soil reaction values are obtained. The aim is to determine the actual coefficient of soil reaction value, by evaluation through a model, including the dynamic effects through a more realistic approach of these flexibility values, obtained experimentally. By this way taking into account the dynamic effects in the iiuinerical model the flexibil515
ity values are drawn by giving different coefficient of soil reaction values. The coefficient of soil reaction value of the soil, which is the coefficient of soil reaction value, used in the curve, drawn in the numerical model, with which curve the flexibility values fit, obtained in the experiment. Through a prepared computer program and in the cases \Then the pile is embedded or articulated and at all kind of soil terms the results of the experiments, made by a vibrating apparatus, may be evaluated. Thus a more realistic coefficient of soil reaction value is obtained. REFERENCES Acar. M.11. 1953. Analyse theorique et experimentale du coinporteinent des pieux. These de docteur-ingenieur, Universite Pierre et Mark Curie, Paris. Baguelin, F. & Jezequel J.F. 1972. Etude expirimentale du comporrement des pieux sollicites horizontalement, Annu1e.s I. T.8.1’.P.11‘’ 297, Paris. Bowles, J.E. I9SS. Foundation Analysis and Design, Mc GrawH i 1I, Ne\v-Ynrk. Rriard, A. 1978. Mesure du coefficieiit de raideur horizontale d‘un pie11 ancre dans le sol a I’aide d’un excicateur a Balourd, Contrnt DBTPCICEBTP, Paris. Chtti, L. & PoLtlos I l.G. 1993. Analysis of pile-soil interaction iirider lcrci~al iociding using infinite and finite elements, i 'amp m d Gcorc‘chnics,Oxford. England.
516
Application of Stress-Wave Theory to Piles, Niyama & Beim (eds) 02000 Salkema, Rotterdam, ISBN 905809 150 3
Analysis of crosswise vibration of pile Chen Renpeng, Zhu Bin & Chen Yunming Geotechnicul Engineering Institute, Zhejiang University,Hmgzhou, People's Republic of Chirzu
ABSTRAT: A rational mechanical model of foundation of precast concrete pile on layered eiastic base when driving is established in this paper. Through dynamic anal!isis. the nurneric solurion of forced reaction of the foundation of single pile which eccentric dynamic \erticill load XIS on pile top is provided. And effects of excitation intensity. impact duration of pile lmiimer, diamcter or length of side of pile. eccentricit), of pile hammer, foundation modulus and coefficient of viscosity of soil to the dynamic response of pile are analyzed. The primary analytic method of studying the pile response LI hich intense axial excitation and moment excitation act on. which is very important when studying the vibration of excited precast concrete pile. pile, and thcq haic not farther discuseJ thc dynamic response of tlie pile ~ h c heccenrric dynamic vertical load acts on. 'Thc niomcnt of pile h m n i c r ncting on pile top is an intciisi'c7c impact loxi whc11 Y , , 2 . . Y , 7
'
11-2
,
Yl ) I - ,
>
I'
Y,,,,
(7)
the column vector of load in center of every pile unit is:
the program based on the method mentioned above is provided in this paper, which can present the displaceinent response of every pile unit at any time, and so do the shear force and the moment. (there is no algebra function of in and c in equation (4) and ( 5 ) of the pile unit above gro u 11 d ) where h = step length of center different H = ( H=pile length), s = step length of time, n-3
3 EFFECTS OF CROSSWISE VIBRATION OF PILE When pile haininer impacting on pile top, the pile-soil system is vibrated by the excitation of pile hammer to pile, and the vibration is transient and nonperiodic. Commonly, the time of energy translation is short than cycle period of auto-oscillation of the pile, but the duration of response can be several times of cycle period of auto-oscillation of the pile. There are 5 main factors effecting on the crosswise vibration of pile, which are excitation intensity, impact duration of pile hammer, eccentricity of pile hammer, foundation modulus and coefficient of
a , 6 is constants: a 3 0.25(0.5 + 6)' , 6 3 0 . 5 , s, =(j-2).h,
I, =i..r,
hi, = N o .sinat, ; when 1,
when 0
- F,,,I- F,) / v Calculate mean damping factor C,,,,. for above range Now static resistance Fu can be calculared at all points
E, = F.m - Cmw, .V- Fd D r w static load-displacement diagram with F,, and
562
U
Application of Stress-Wave Theory to Piles, Niyama & Beim (eds) 02000 Balkema, Rotterdam, ISBN 90 5809 1503
ynote lecture: Three-dimensional finite element analysis of statnarnic load test Tirawat Boonyatee & Makoto Kimura Kyoto University,Japan
Feng Zhang GifLl Universiq,Japan
ABSTRACT In order to stud! the inechaiiisin of pile-soil interaction during statnamic loading. a FEM program is developed Since it n i l 1 also be applied to the statnamic test of piles in lateral direction. batter piles. group pile. and pile raft foundation, the anal! sis S J stem is formulated i n three-dimensional space Elasto-plastic model is used for the ground material Joint element is also used for slip at the pile-soil interface The analysis program uses the direct integration scheine to solve the elasto-plastic anal!sis in time domain Laboratory tests of single piles are analjzed to iiivestigate the validit! of the developed program Firstly. the simulatioiis of static load test are done and compared \i7ith the nieasurcd data from the tests Then. the simulatioiis of piles uiider statnamic are conducted uiider the same grouiid coiidition Finall!. thc comparisons betu cen the static and statnamic responses of piles are discussed 1 INTRODUCTION
At present, the primary approach to investigating pile behavior under statnamic loading is to make a comparison between the results of full-scale statnamic and static load tests. Unfortunately, this calibration method has an inherent problem in its testing consistency, as was addressed by Amir & Amir (1995). Under the uncertainty and complexity of the target ground, the situation whereby any model can be validated by calibrating statnamic load tests against corresponding static load tests can be demonstrated as follows: 1 If a comparison is made of the results of tests on the same pile, the tests have to be done in sequence. Therefore, the quality of the later tests invariably deteriorates due to the induced residual stresses from the former tests. 2 If a comparison is made between the results of tests on different piles, significant samples should be tested in order to achieve some degree of reliability over the ground uncertainty.
lnstead of studying the data from the field tests, the mechanism of the piles during statnamic loading can be thoroughly investigated by the numerical analysis approaches. Applying the fact that the system of interest is axis-symmetry, Matsumoto (1998) applied a 2D-FEM program to analyze the behavior of a single pile under vertical loading. Although a twodimensional analysis is simpler and the solution can be found faster than with a three dimensional analy-
sis, it cannot be applied to a complex system such as a pile-raft foundation or batter piles Moreover, it is obvious for a horizontal loading case that the analysis should be done in three-dimensional space In order to develop a unified tool that can be applied for various types of problems, a threedimensional finite element analysis program called DYSC (Dynamic and Static systems analysis Code) is originally developed No-tension criteria and a simple elasto-plastic model based on the DruckerPrager theory are applied as the yield functions for the model ground An interface layer element is inserted at the interface between the pile and the soil For the pile body, a linear elastic relation is used since the applied force is lower than its yielding point It is thought that before applying the DYSC to the simulation of piles, its fidelity should be checked by a physical model For calibration purposes, a comparison of the analytical results to the laboratory test results is preferred to the results from the field tests The reasons are due to the complexity of the ground and the qualities of the geotechnical parameters required for the analysis In our previous study, a small-scale statnamic loading device (3SLD-Mkl) was used to conduct laboratory tests of piles under vertical loading (Kimura et a1 1998) This device is used to conduct experiments under a controllable environment Consequently, guaranteeing that the series of tests are done under the same conditions is possible Accompanied with the data from those experiments, a comparison 563
between the simulation and the experiment is made and reported in this paper. 2 DETAILS OF ANALYSIS
2.1 Inyzrt yaranieters To determine the necessary parameters used in the simulations of the statnamic load test, trial calculations are made. As shown in Figure 1, an analysis is conducted in the half area of the pile-soil system. Young's modulus of the pile is back calculated from the flexural test results. Poisson's ratios for the pile and the soil material are assumed to be 0.2 and 0.333, respectively. A parametric study has been made to determine the Young's modulus of the model ground. From various trial calculations and comparisons with data from static load tests, a Young's modulus of 4.9 MPa and a frictional angle of 32 are selected as rational quantities for the ground materials. The frictional angle of 32 for soil gives the calculated ultimate load at more or less the same level as the measured data. A Young's modulus of 4.9 MPa controls the shape of the loaddisplacement of the pile before failure. Using these values, a satisfactory approximation for pile response can be obtained. The load-displacement plot from calculations and measured data is shown in Figure 2.
0
1
2
3
3
5
6
D i sp1acein ent (inin) Figure 2. Load-displacement relations from simulation and experiment.
Table 1. Prouerties of materials Sand
Pile
Density (kg/m3)
1467
2150
Young's modulus (MPa)
4.90
5.00s103
Poisson's ratio
0.333
0.200
Friction angle ( )
32
Dilatmcy angle ()
9
The properties of each material determined from an inverse analysis are summarized in Table 1. 2.2 Pile-soil inteiface model
The Mohr-Coulomb yield criterion is used for the interaction between the pile and the soil. It states that failure will take place if the magnitude of shear stress (z) on the failure plane is equal to the value given by the following relationship:
in which 1 1 denotes the absolute value, 0, is the normal stress on the failure plane, and $0 and CO are material constants for the pile-soil interface. In this study, adhesion, or CO, is assumed to be zero. The frictional angle of the pile-soil interface is assumed to be 0.9 times the frictional angle of sand. Equation (1) can be written in the form of yield fbnction F as
F = jzlFigure 1. Analysis mesh.
564
CT,
tan$o
If the material is sheared to the yield surface and the associated flow rule is adopted, the rates of plastic normal strain d&; and shear strain 7' are given by
which implies 0.0
(4) Increments in shear displacement along the plane are accompanied by increments in normal displacement. The dilation of the shear plane will go unbound under yielding. To avoid this unfavorable behavior, the non-associated flow rule is adopted for the pile-soil interface. By introducing dilatancy angle y, the plastic potential function can be written as
In this study, no dilation is assumed for the pile-soil interface, i.e., w-+ 0.
3 SIMUL,ATION ON STATNAMIC LOAD TEST
After necessary parameters were back calculated from the previous section, a dynamic FEM analysis of a pile under Statnamic loading is done and compared with the experiment results. The loading rate dependency of the ground response is represented by a constant damping parameter. In this simulation, velocity and acceleration are calculated from the displacement by the Newmark method. The stiffness matrix (IQ the damping matrix (C), and the mass matrix (A4) are calculated from the following equations:
K = JB'DBdd 1.01
(7)
565
0.2
0.4
0.6
0.8
1.0
1.2
Di sp1acein eiit (m in ) Figure 3. Load-displacement relations from statiiaiiuc simulation and experiment.
where N is a so-called shape function or displacement interpolation function, R denotes a displacement to the strain transformation fbnction, and 11 is a strain to stress-strain transformation function. Parameter ,U is used to represent the damping constant per volume of interested material in the same manner as the density (p) is applied in the mass matrix. Based on trial calculations, a damping constant of 3.9 MN/(m/s)/m3 (,U as defined in Equation (7)) is used as the material constant for statnamic simulations. The estimated load-displacement relation and the test data are shown in Figure 3 . Note that the initial displacement in statnamic load tests does not conform to that in the static load tests. The initial displacement in the statnamic load tests is a little bit smaller than the corresponding value in the Static load tests. This may contribute to the loading rate effect during the equipment installation process. To correct this inconsistency, the load-displacement relation (from the experiment) of the statnamic load test is shifted to the right in order to match the 'true' displacement in the static load tests. 4 COMPARISON BETWEEN COMPUTED STATIC AND STATNAMIC LOADING RESULTS
The distributions of displacement along a pile under static loading are shown by Figure 4. At the peak load, the elastic contraction represented by the difference in settlement between the pile head and the toe is about 0.05 mm. When compared to the overall settlement, this contraction is about 5% of the pile head settlement. The distribution of axial force along the pile is shown in Figure 5 . The application of load is sustained almost totally by friction resistance, with only a small amount of force being transmitted to the pile tip. The ratio of the end bearing resistance to the shaft resistance is about 1 :4.5.
St3t.k
150 N +250 N +400N 450 N
+
s tatnarfl k: - 0 - 150 N - 0 - 250 1\1 - 0 -400N
- H - 450 N
~"'"1'""1"""'"I""'I""'I''"'I'""I'
()
0
02
0 -4
00
08
10
D 1s p lacem en t ( m m) Figure 4 Distnbdion of displaccnient along pile under static loading
0
50 100 150 200 250 300 3% -100
Load (N) Figurc 5. Distribution of axial forcc along pilc undcr slalic loading.
The pile velocitv versus time relation is shown in Figure 6. At the maximum displacement, the velocity of the pile head is equal to zero. At the same time, the velocity of the pile tip is almost identical to that of the pile head. This implies that the pile behaves as a rigid body at the unloading point. Axial force distributions of the pile during statnamic loading as well as a comparison with those of the static load test are shown in Figure 7. At the same applied load level, the axial force distributions of the two cases are almost identical. There are no substantial changes in load sustained by the pile tip. Increases in capacity mainly contributed to the shaft resistance. The effect of a stress wave is not observed in this analysis. 5 CONCLUSIOK
Figurc 6 Vanation
III
When compared to field tests that have the inherent problem of ground uncertainty, the model tests presented here show the possibility for conducting Statnamic load tests under one uniform condition. Only by this approach can the data be interpreted on absolutely the same basis. Based on the results presented here, the following conclusions can be summarized: 1 The finite element analysis (DYSC) shows that the elastic contraction of the pile is relatively small. This value represents about 5% of the total displacement. From the load distribution plot, the applied load is mainly supported by shaft friction. The model pile is thought to be a friction pile.
vclocitj of pilc licad i. From the test data the force-time curves give the most information about the behavior of the testing apparatus. Jumping heights were recorded, but are not presented here. Figure 8, shows the force-time curve for one of the first 4MN tests on the 600 mm piles. Figure 9. shows a typical force-time curve for a conventional Statnamic testing using concrete and steel masses. The similarity between the two graphs shows the effectiveness of the water reaction mass assembly. Closer examination of Figure 8 shows a loading event of just over 100 ms duration. Figure 9 is at least 10 ms shorter in duration. This was the expected result since the land test was performed with a total reaction mass of 22,000 kg, whereas the water test used closer to 25,000 kg. Although the results were not surprising they do offer direct evidence of the effectiveness of the contained water as reaction mass for the Statnamic test. Also of ncte was the successful operation of the mechanical catching mechanism. As the Statnamic silencer reached the high-point of its jump it was noted that the subsequent downward movement was abrupt - similar to the 1998 experimental testing
Figure 9 Load vs Time using concrete and steel as reaction
when the catching cylinders were deactivated. As a consequence, the descending silencer and reaction masses produced a sharp impact load on the catching frame. Several of the steel teeth as well as the mechanical latches showed noticeable wear. The catching device functioned well nonetheless, and proved the idea of a mechanical catching mechanism to be sound. 4 SUMMARY AND CONCLUSIONS
The main conclusion drawn from the work described in this paper was that water could be used successfully as reaction mass for Statnamic testing. The theoretical work by Middendorp and Courage was useful in the design of the water containers. In particular, the use of a valve at the base of the water containers was highlighted in their work as an important element. In the 1998 experimental testing the mass of the steel containers was high in relation to the mass of contained water, hence the Middendorp and Courage valves had only a moderate effect. The containers used in the Lake Charles contract testing omitted the valves entirely. As the percentage of contained water increases in relation to the overall reaction mass (this is likely for larger capacity tests), it is theorized that the valves will become increasingly important. The experimental work of 1998 was the first testing of its kind, and proved the general concept of using water as reaction mass. Pressure transducers contained in the submerged reaction masses indicated that the drag forces of the container provided an additional resistance to the Statnamic event of 3030%. The moderate jumping heights observed in this testing also demonstrated the usefulness of drag forces. The final proof of the concept of using water as reaction mass was the contract testing in Lake Charles, Louisiana. This testing proved that a lowcapacity mechanical catching mechanism could be used to catch the water reaction masses. It also
Figure 8 Load vs Time usin$ water as reaction mass
615
proved that such a test could be assembled freestanding on the test pile, without the need for a support frame or special support structure. The success of the 1998 experiments as well as the Lake Charles testing point toward the future use of water as reaction mass for Statnamic testing in marine environments.
REFERENCES Applied Foundation Testing, Inc. Phase I1 - Report of Statnamic and Dynamic Load Testing, New Ship Berth at Contraband Bayou - Contract “A“. Prepared for The Lake Charles Harbor and Terminal District. c/o CBK Soils Engineering. Inc.. September 14, 1999. Applied Foundation Testing. Inc. Phase I - Report of Static, Statnarnic and Dynamic Load Testing, New Ship Berth at Contraband Bayou - Transit Shed. Prepared for The Lake Charles Harbor and Terminal District, c/o CBK Soils Engineering. Inc.. May 28. 1999. Baddour. R., 1998. Calculations modeling the reaction mass of hydraulic dynamic effect: Report fiom the tiniversigi of Western Ontario, Depr. of Civil Engineering, London, Ontario, Canada, I998 Janes. M.C. 1998. Statnamic testing using submerged reaction mass, Second International Starnamic Seminar; preliminary proceedings, Tokyo, October 29-3 I 1998. Tokyo: Japanese Geotechnical Society. Jonker. G. 1999. Use of a ballistic hammer for installation of anchor piles for off-shore structures. Conrinuous advances in Moorings Ce Anchors; proceedings, Aberdeen, 26-2 7 M ~1999. J London: IBC UK Conferences Limited. Middendorp. P.. and W.M. Courage, “STN above-water testing with undenvater containers“, Reporr ,from TNO Building and Construction Research, Delfi, Netherlands, 1995.
616
Application of Stress-Wave Theory to Piles, Niyama & Beim (eds) 02000 Balkema, Rotterdam, ISBN 90 5809 150 3
Introducing statnamic load testing in Europe: Case studies in the Netherlands G.J. J.van Ginneken & l? Middendorp irTV0 Profound, Delft, Netherlands
ABSTRACT: A short introduction is given of the principles of the statnamic (STN) pile load testing. Examples are given of the applications of STN in The Netherlands and some European initiatives. Practical experiences with the use of the 4MN STN device with catching mechanism are described. Concluding economic considerations will be given when applying STN in practice.
1 INTRODUCTION
The duration and loading rate of a STN test can be controlled with the volume of the burning chamber and the shape of the cylinder and piston, the amount and type of fuel and the amount of reaction mass launched. As a result, the load can be introduced more gradually and for a much longer dura-
Statnamic started in the early nineties to satisfy construction Industry’s demand for cost effective and accurate means of testing high capacity foundations. Berminghammer Foundation Equipment of Canada joined with TNO Building and Construction Research of The Netherlands to develop STN. At this moment over one thousand STN load tests have been performed in Canada, United States, Japan, Malaysia, United Kingdom, The Netherlands and many other countries in the world. The equipment available at present can perform STN load tests starting from 0.1 MN up to 30 MN. Design and construction of even higher capacity STN devices (60 MN or more) can be expected in the coming years. The principle of STN is based on the launching of a reaction mass from the pile head. Launching takes place by generating high pressures in a cylinder, caused by the burning of a special fuel. As a reaction on the launching the pile is gently pushed into the soil. The load exerted on the pile head is measured by means of a load cell. The displacement of the pile head is registered by means of a special developed laser sensor. Load cell and laser sensor are integrated components of the STN loading device. No instrumentation has to be installed on the pile shaft. The required reaction mass is 5% of the force generated. The high capacity STN devices are not restricted to test single piles, but also allow the testing of pile groups and structural elements such as bridge piers and spread footings. Because the principle of STN is based on the acceleration of masses, piles can be tested in any direction, also horizontal and under batter.
Fig. 1. Force versus time diagrams
617
tests indeed showed the same or similar results as static tests.
tion compared to dynamic load testing. As an example the force time diagrams of dynamic load test, a STN load test and a static load test have been presented in Fig. I . The long duration of the STN loading keeps the pile under constant pressure and tension stresses cannot develop. The central location of the STN loading device on the pile top guaranties an axial introduction of the load onto the pile. The long duration of the STN loading causes for all pile levels a similar displacement behaviour as can be observed with static load testing. This justifies a simple modelling of pile and soil, in which stress wave phenomena do not have to be taken into account. The pile is considered to be a mass on which the STN force, the inertia force and soil resistance are acting (Fig. 2).
During the first years the desire to use STN for higher loads quickly rose. Statnamic devices to measure up to 30 MN were developed and are still used in larger projects. From the beginning the economic gains in these larger projects were evident. After first setting ground in Asia in mega construction projects STN is now quickly developing a wide acceptance in the USA as well. Testing is steadily growing, having performed up to present some 300 STN tests in the USA alone. The way STN was introduced in the USA was identical to its acceptance in the Far East. First comparative studies were carried out between static and STN load tests to confirm the reliability and accuracy of STN before the transportation authorities adopted the method. Several case histories in the Netherlands and Germany during the early nineties showed satisfactory and good agreement between static load behaviour from STN results and static load test results. It was concluded that: During STN load testing, pile behaviour can be modelled as a mass on which inertia forces and soil resistance are acting. This allows the simple calculation of the static load behaviour. Statnamic load testing can be performed on piles situated in soils with a strong dynamic response. The point of unloading (maximum displacement) in the STN load displacement diagram allows the direct calculation of the maximum static soil resistance during testing. Strain rate sensitivity has to be included for clayley soils.
Fig. 2. Forces acting on the pile during STN loading 2 HISTORY OF STATNAMIC DEVELOPMENTS
Worldwide during the last decade over one thousand statnamic tests have been carried out. The first developments of STN started in Canada, USA and the Far East where alternative ways of pile testing were introduced to accommodate for the high demand for pile testing in strong developing economies. There was much to gain, especially in the larger construction works where pile testing appeared to be cumbersome and expensive. Numerous comparisons between static and STN tests on driven and cast-in-situ pile have been made to convince authorities that STN proves to be good alternative for static load tests. Acceptance was acquired when authorities had seen that the results obtained from the STN
Europe has been somewhat reluctant to accept STN as the alternative testing method. Reasons for this are the existing building codes based on a traditionally strong preference for static load testing for well defined and safe cases. Moreover, for driven piles over the years a market for dynamic load testing had been developed and is now well established. The incentive to adopt alternative methods has not been large because existing regulations did and in many case still do not mention STN as an acceptable method for pile load testing and the market has the perception that STN testing is expensive. The question whether to perform pile testing or not in many cases depends on the funds available and the judgement of the experts on site. Money appears to 618
be dominant over quality assurance and as a result often a lot of construction works go without regular testing of foundation piles. The reason is that static load tests are considered too expensive, time consuming and therefore cumbersome. Dynamic load testing is not always recommended because of its limitations.
model for the strain rate sensitivity of clayley soils under STN loading conditions. Demonstrations have been trend-setting in adopting this new technology into the market. The method used by TNO Profound to introduce this new concept is to organise seminars and training courses for users and potential users to acquaint themselves with this new technology. This introduction involves a theoretical part in which the geo-technical background of STN is being explained and a practical part in which the actual testing is demonstrated.
For STN to be attractive applications are to be found in the lower pile loads testing range up to 2.53.5 MN and competition with established low cost static and dynamic testing methods is strong. In addition, regulations in most European countries are not yet adjusted in considering STN as an acceptable means for pile Ioad testing replacing or in conjunction with static load tests.
Collaboration exists with the leading Universities and Research Institutes in the countries involved securing adequate technical support, acceptance and being able to open steps towards introduction of STN into the national regulations. This means training, educating and continuously introducing technicians into the world of STN.
3 INTRODUCING STATNAMIC IN THE NETHERLANDS AND EUROPE
New countries covered are Spain and Turkey, where they are very interested to get STN introduced. Preparations are made for a first STN demonstration project in Barcelona. In Poland and Hungary progress is made to get STN into the regulations and perform static-STN demonstrations in 2000. Infrastructural EU supported investments for the preaccession countries are the driving force for the introduction of STN.
In response to the increased interest in statnamic in Europe TNO has decided to set up TNO Profound (Professional Foundation Diagnostics) responsible to introduce STN further in Europe. In combination with the knowledge base in foundation technology present at TNO Building and Construction Research a strong basis has made to make STN a success following the examples in the Far East and the USA. 3.1 Stntnarnic load testing with a 4MN device From its start in March 1999, TNO Profound has a 4 MN STN device at its disposal at a central location in The Netherlands for testing on the European continent. This has opened up the market for STN in Europe. Steps are taken to position a second 16 MN device in the Central European market with the aim to have a STN testing facility available on local markets to conduct testing services at commercially attractive prices. In 1999 TNO Profound has been engaged in several research projects in Europe in which the results of Static and Dynamic Load Testing have been compared with STN test results on the same and/or similar piles. Projects have been carried out with Jacbo (The Netherlands), WTCB (Belgium), the Technical University of Budapest (Hungary). Preparations are being made to initiate similar projects in Germany, although Germany has a strong tradition in static load testing the willingness to apply STN in Germany as well is rising now STN tests become economically more attractive. TNO Building and Construction Research is participating in a research project with the University of Sheffield to develop a
3.2 Major considerations for the introduction of Statnanzic on the European continent Regulations The tradition in most European countries is to conduct static load tests on specially made test piles. Loads to be applied range between 1.5 - 2.0 times the design load. These static load tests require a lot of preparation work and piles are to be carefully selected and prepared for testing. This is a costly and cumbersome activity. For this reason static load tests are virtually abandoned in The Netherlands for proof testing and are becoming less popular in other countries counties like Belgium as well. Time and construction costs are increasingly under pressure squeezing cumbersome testing practices to the bare minimum. In Europe most countries the use of alternative load testing methods is not foreseen in the national regulations. This hampers the wider introduction and acceptance of other innovative and more convenient methods. Mainly in Germany dynamic load testing has gained interest because of its convenience to use. A pile driving hammer or drop hammer can be used. Piles to be tested can be selected randomly. The same counts for STN where the reaction mass re-
619
quired is only 5% of maximum load to be applied. Construction engineers of local organisations may decide to accept DLT or STN when judged sufficiently suitable for its purpose. Their application depends on the availability of a testing capacity and is mostly used on a case-by-case basis, where convenience, time pressure and economics play a major role. For cast-in-situ piles STN is the preferred method (P.Middendorp, 2000). TNO Profound is promoting to consider also alternative testing methods like STN into the national regulations of EU countries. This is believed to be an essential prerequisite to get STN widely accepted in Europe. STN demonstrations with the involvement of the academia and regulation authorities are required to get decision makers acquainted with the method.
Statrzanzic as the only feasible alternative for testiiig
The major applications of STN in The Netherlands have been cases where the alternative testing methods where not seriously considered because of technical limitations (DLT on cast-in-situ piles would not generate satisfactory results), time pressure or undesired excessive costs. STN was not requested to replace routine static load testing, but was considered to identify the load bearing capacity of existing piles for re-use or in situations of doubt about the performance of disturbed soil profiles where conventional calculation and testing methods could not be applied. Groundwater seepage has been one of the major causes. STN would also be ideal in situations where a lot of piles are to be tested in a short time frame. For such larger projects STN is still lacking sufficient EU based reference projects.
Ecorzoinic considerations The price per pile for static testing varies considerably depending on the load to be applied and the specific site situation in which the tests are performed. When more piles are to be tested economies of scale cannot be made since each pile needs to be individually prepared. When one compares static load testing with STN, the testing costs of STN are usually 30-60% lower. The real economic gain is not so much in the testing costs as such, but in the considerable time saving when applying STN. For STN the preparation time in minimal, no reaction piles are required, piles for testing can be selected randomly and testing results are immediately available. This allows for instant decisions on additional testing requirements. For calibration purposes of STN it is always recommended to conduct one comparison test with a static load test in unknown soil conditions.
The price of STN and dynamic load testing is largely influenced by the costs of mobilizing the testing equipment. When the testing equipment is used for testing of one pile only the price remains relatively high. When more piles are tested on one project site the price per pile drops considerably. When comparing STN with DLT, pile preparation for STN is easier, but more time is required for the installation of the testing device. Four to 6 piles can be tested a day with a STN device with catching mechanism. Transportation costs in Europe are considerably higher than in the USA and allowable transportation weights per truck are maximized. Therefore in Europe it is preferred to move empty containers to be filled on the site with load material, whereas in the USA it’s easier to transport concrete/steel reaction masses. 4 CASE STUDIES OF STATNAMIC IN THE NETHERLANDS
4.1 Statizamic pile testing in Rotterdam Project Data Location : Rotterdam Harbour : Wilton-Feijenoord Time : June 1999 Client : Van Hattum & Blankevoort Testing : Prefabricated piles 420x420 4 piles, 23 m long (new) 2 piles, 24 m long (existing) Load applied : 2 -4 MN Project Description At one of the piers in the Rotterdam harbour a large construction crane was present for the construction of offshore vessels. It was intended to extend the moving range of the crane to an old reconstructed pier by using the existing foundation piles. It was questioned whether the existing piles under the pier would be suitable to bear this additional load. TNO Profound was requested to apply maximum loads in the order of 3.5 MN to check the load bearing capacity of the piles. Project Characteristics The piles had to be tested over water under difficult site conditions, limited availability of crane capacity and considerable time pressure (sea vessel to be loaded). static load testing and DLT not feasible because of shoreline conditions and accuracy. The 4 MN STN device was moved from pile to pile in one
620
Fig. 4. STN Testing at Eindhoven 4.2 Stutnamic pile testing in Eindhoven
Project Data Location : Eindhoven Railway Station : Large Shopping Complex Time : October 1999 Client : Inpijn-Blokpoel Consultants : Jacbo Avegaar piles (cast-in-situ) Testing 15 piles, 2 1 m long (new) Load applied : 2,5 - 4 MN
Fig.3 .4MN STN Device in Rotterdam. piece. Careful and accurate installation of the STN device above water appeared a time consuming process. Two piledday were tested with one- two testing cycles per pile.
Project Description A large shopping complex was to be built. Construction works were suffering delays and trouble with the foundations would negatively affect the timely progress of works. The piles were located in a 6-7 meter deep construction pit and testing had to be carried out on selected piles. After placing the foundation piles considerable groundwater seepage had occurred and maintained to grow during the construction works. The seepage was stopped through injecting the soil, but the soil stability around the piles was believed to be disturbed. Doubts arose whether the piles under the main structure would perform according the calculations based on soil investigation results and their capacity had to be determined.
Results The following loads were applied to the piles: Max load Displacement (in mm) (static) in kN (total) (permanent) Pile 1: 4190 Pile 2: 3980 Pile 3: 3960 Pile 4 (step 1): 2750 Pile 4 (step 2): 4340
20,s 19,5 18,s 12,l 17,6
395 377 530 2,o 1,5
Project Characteristics First a reference test was carried out on a pile installed in an area which had not been affected by seepage. The results of the STN test on the reference pile were compared with the test results of the seepage affected piles. It was also indicated that the amount of piles to be tested would depend on the testing results. The piles selected for testing were indicated from the ones most severely affected in concentric circles going outwards towards less seepage affected piles.
Conclusion The conclusion was that the existing piles performed above expected calculated bearing load and could be safely used for the pier extension. After the first tests the client wanted the tests to go further up to failure to determine ultimate load capacity. The tests showed that the performance of existing piles may well exceed calculated capacities.
621
Results The aim was to minimize the time spent on testing. During the first tests it appeared that cyclic testing would have a positive influence on the performance of the piles in working conditions. It was decided to conduct 2-3 cyclic loads per pile on the piles most seriously affected. The following loads were applied to selected piles: Max load Displacement in mm (static) in kN (total) (permanent) Pile 1 : 2220 Pile 2: 2130 Pile 3: 3460 Pile 4 (step 1):2450 Pile 4 (step 2): 31 10
11,l
8,1
27,9 9,1 18,5
790 4,l 21,5 2,4 3s
Project Characteristics This project is typical for a common situation in the Netherlands where the standard procedure is not to conduct static load tests on a regular basis, but to rely on the results of CPT tests to approve piling jobs. In case of poor CPT results construction engineers are only inclined to approve the foundations when the piling contractor can prove that the piles installed perform within the limitations. TNO Profound was requested to test and provide an independent judgement on the load bearing capacity of the Omega piles. Cyclic STN tests on a reference pile and on a questionable pile were compared to assess the difference between the two and to judge the acceptability of the tested piles. On both piles 4 cyclic load tests were carried out up to failure capacity to fully satisfy the demands of the construction engineer.
Conclusion It was concluded that the piles proved to be just within the margin of acceptance. Signal matching with TNOSTN revealed that the soil skin friction in the upper 4 m layer was reduced to only 35% to of the total due to the effect of seepage. The total pile capacity was seriously affected by this, but the lower soil layers compensated for the loss occurred and would still have sufficient capacity to provide the necessary support for these pile foundations.
4.3 Statnamic pile testiizg in Utrecht Project Data Location : Utrecht : Jaarbeurs Fair Complex : December 1999 Time Client : Jacbo/HBG : Jacbo Omega piles (Auger) Testing 2 piles, 18 m long (new) Load applied : 2,5 - 3,5 MN
Fig 5. 4MN Statnamic device in Utrecht Results The aim was to minimize the costs and therefore the amount of time spent on testing. The results were satisfactory and the foundation piles were accepted by the construction engineer. The following loads were applied to the piles:
Project Description At the Jaarbeurs fair grounds of the Municipality of Utrecht, new high rise office buildings were planned for construction. For the Omega pile type it was a routine procedure to compare the CPT results derived from tests before and after the pile construction works. These values are then being compared to check for a reduction of CPT values caused by the installation of the piles. In one of the a corners of the building site the CPT results after piling showed values lower than the set limits. STN tests were carried out to verify that the capacity and settlements of the piles would be within the allowable limits.
Max load
Displacement in mm (static) in kN (total) (permanent)
pile 1 pile 2 pile 1 pile 2 pile 1 pile 2 step 1 step 2 step 3 step4
622
1690 2170 2190 3110
1850 16,3 19,2 2570 16,O 15,5 2730 9,9 36,l 3140 18,5 38,5
12,O
13,O 6,O 9,O
10,O 12,O 28,O 31,O
Conclusion Statnamic proved to be a simple and effective way to test the performance of cast-in-situ foundation piles in cases where CPT tests would give doubtful results. STN testing confirmed that in some cases the interpretation of CPT tests alone could lead to questionable conclusions. With STN an effective tool was provided to verify independently the actual performance of the pile tested.
5 CONCLUDING REMARKS When considering the STN testing worldwide it can be concluded that nowadays every other day a STN load test is carried out somewhere in the world. USA, Asia and Japan are taking the lead. Europe is a region where STN has a large potential to grow. TNO Profound has actively started to promote STN in Europe and aims to motivate centres of expertise in European countries to initiate STN activities. STN can replace static load testing, however, when deriving design rules, static load testing will remain the preferred testing method.
6 REFERENCES Bermingham P., Janes M., 1989, An innovative approach to load testing of high capacity piles, Proceedings of the International Conference on Piling and Deep Foundations, London, p.409-413. Middendorp, P., Bermingham P., Kuiper B., 1992, Statnamic load testing of foundation piles. 4th International Conference on Stress Waves, The Hague, Balkema Middendorp, P, 1993, First Experiences with Statnamic Load Testing of Foundation piles in Europe, Proceedings 2nd International geotechnical seminar on Deep Foundations on Bored and Auger Piles, Gent, p .265-272, Balkema Brown, D.A., 1994, Evaluation of Static Capacity of Deep Foundations from Statnamic Testing. Geotechnical Testing Journal, Vol 17, No.4, American Society for Testing and Materials Matsumoto, T., Tsuzuki, M., 1994, Statnamic Tests on Steel Pipe Piles Driven in a Soft Rock. International Conference on Design and Construction of Deep Foundations, Orlando, U.S. Federal Highway Administration Middendorp, P, Bielefeld, M.W., 1995, Statnamic Load Testing and the Influence of Stress Wave Phenomena, First International Statnamic Seminar, Vancouver 623
Middendorp, P., Foeken, R.J. van, 1998, When to Apply Dynamic Load Testing and Statnamic Testing, 2nd Statnamic Seminar, Tokyo Middendorp, P, Ginneken, G.J.J. van, Foeken, R.J. van, 2000, The advantages and disadvantages of D namic Load Testing and Statnamic Load Testing. 6' International Conference on the Application of Stress Wave Theory to Piles, Sao Paulo, Brazil
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Application of Stress-Wave Theory to Piles, Niyama & Beim (eds) 02000 Balkema, Rotterdam, ISBN 90 5809 750 3
The advantages and disadvantages of dynamic load testing and statnamic load testing l? Middendorp & G.J. J.van Ginneken TNO Profound, Delft, Netherlands
R. J.van Foeken TNO Building and Construction Research, Department of Structural Dynamics, Rijswijk, Netherlands
ABSTRACT: Pile capacity testing by high strain dynamic loading methods is widely applied because of its economy and efficiency compared to static load testing methods (SLT). Frequently applied dynamic loading methods are dynamic load testing (DLT) and statnamic testing (STN). The paper will deal with the very often raised question in practice: What are the advantages and disadvantages and when to apply DLT and STN when pile type and soil conditions are known. Special attention is given to DLT on cast in situ piles, and complicating factors like limited knowledge of concrete material properties and pile shape. The suitability of DLT and STN is discussed for cast in situ piles and driven pre-cast piles by the evaluation of reliability, economy, mobilization of capacity and the chance on pile damage.
1 INTRODUCTION
Pile capacity testing by high strain dynamic loading methods is widely applied because of its economy and efficiency compared to static load testing methods (SLT). The most popular dynamic loading methods are dynamic load testing (DLT) by an impact hammer and statnamic testing (STN) by launching a reaction mass from the pile head. DLT introduces a short duration shock pulse into the pile. STN generates a relative long duration push load onto the pile head. Extensive descriptions of load testing methods and comparisons are published by Holeyman (1992) and Karkee et a1 (1997). However these papers do not deal with the very often raised question from practice: What are the advantages and disadvantages and when to apply DLT and STN when pile type and soil conditions are known. The answer to this question will be treated in the next paragraphs . Special attention is given to DLT on cast in situ piles, because the calculation of the pile load is based on signals from strain transducers mounted on the pile shaft. So for DLT the pile load calculation depends strongly on pile material and cross section properties and factors complicating the analysis like limited knowledge of concrete material properties and pile shape are discussed.
Fig. 1 statnamic test on a cast in situ pile
625
-
-
economy mobilization of capacity chance on pile damage
2 THE APPLICATION OF DLT AND STN ON CAST IN SITU PILES For cast in situ piles both DLT and STN are performed a certain period after pile production, to allow the piles to reach the required compressive strength to withstand the test loads. For DLT strain and acceleration transducers are mounted on the pile shaft near the pile head. The load displacement behavior is calculated by signal matching. For STN the load displacement behavior is calculated in most cases by the Unloading Point Method (UPM), however signal matching techniques are also applied.
2.1 Accuracy in load measurement for STN With STN the load is accurately measured by a calibrated load cell placed on the pile head. The measured load is not dependent on the pile properties. The load measurement error is less than 0.1% of the maximum range of the load cell. 2.2 Material properties and accuracy in load measurement for DLT
Fig, 2. Dynamic load test on a cast in situ pile.
With dynamic load testing strain transducers are mounted on the shaft near the pile head. The load (F) on the pile head is calculated by multiplying the measured strain (E) with the modulus
Fig. 3. Statnamic piston with built in load cell and laser displacement sensor placed on a cast in situ pile
Finally the suitability of both DLT and STN will be evaluated by taking into account the following points: -
-
Fig. 4. Strain transducer mounted on the shaft of a cast in situ pile
accuracy of the load measurements reliability
626
of elasticity (E) of the concrete and the pile cross section (A). F=E. A.&
(1)
The accurate determination of the properties E and A for bored piles is difficult in many cases. To calculate the force from the measured strain in a pile during DLT we need to know the cross section and the modulus of elasticity of the concrete at the measuring level. For piles with homogeneous material the stress wave velocity (c) is used to calculate the E-modulus with E = c2. p c = 2L/T
(3)
Knowing or estimating the stress wave velocity c we can calculate the pile load at the measuring level with the formula
F= c’.p.A.
(4)
So the derived stress wave velocity has a strong influence on the value of the load measured in the pile. An error in the measured load will result in an error for the pile capacity prediction.
The stress wave velocity is calculated from the time (T) it takes for a stress wave to travel over the pile length (L) from the pile head to the pile toe and back to the pile head (Fig. 5). For this method it is required that the reflection coming from the pile toe is clearly visible in the signals. In Fig. 6 the force and velocity times impedance signals of two dynamic load test are presented. The first case shows a clear toe reflection and the stress wave velocity can be calculated accurately. If the toe reflection is not visible one has to estimate the toe reflection time. However an error in the estimated toe reflection time (T) and stress wave velocity (c) will result in a considerable error in the calculation of the E-modulus. For example a 5% error in the stress wave velocity will result in a 10% error for the E-modulus and corresponding load in the pile. Another option in this case is to rely on an estimate for the E-modulus from the pile material properties. Making an estimate on the E-modulus is difficult because it is not a constant value but depends on the age and the quality of the concrete (Franklin, 1971)(Fig. 7), the loading rate (Sparks et all, 1973 ), and even the temperature of the concrete (Abbasi 1990). For example, for static load testing the modulus of elasticity for concrete is in the range of 28 GPa to 32 GPa while for dynamic load testing it is in the range of 32 GPa to 52 GPa
Fig.5. Calculation of stress wave velocity c from toe reflection
627
Fig. 6 Well visible and no visible toe reflection.
Fig. 7 Relations between dynamic modulus of elasticity and age for concretes made with various aggregates
Another complicatjng factor in determing the stress wave velocity c tor cast in situ piles is the fact that the concrete is not homogeneous. The concrete quality will vary over the cross section and over the pile axis. The concrete in contact with the soil will be of lesser quality than the concrete in the center of the pile and the shaft area that has been in contact with the soil might be the location where the strain transducers are mounted. The concrete quality difference over the pile length is caused by the pouring procedure and the difference in concrete pressure
628
during construction. The quality of the concrete near the toe will in general be better than the quality of the concrete near the pile head. This also means that the stress wave velocity will vary with the pile length. So the stress wave velocity calculated with c=2L/T is a mean value for the whole pile. The modulus of elasticity calculated from the toe reflection represents a mean value for the pile and there can be a considerable difference with the modulus of elasticity at measuring level.
urement results. During STN the load duration is long enough that all pile parts move in the same velocity range. Under these conditions the pile can be considered to act as one mass with a pile rigidity behavior similar to static load testing (Middendorp 1995 ) For this reason pile behavior during STN is closer to static load testing than DLT.
2.5 Economy
Fig 8. Cast in situ pile with bulb.
2.3 Injluence of pile cross section variations on DLT capacity prediction To predict capacity from DLT results, signal matching techniques are the most frequent applied methods, (TNOWAVE, CAPWAPTM).Based on a wave equation computer program calculated signals are matched with measured signals by adjusting the computer soil model and pile model in an iterative way. When signals match it is assumed that the computer soil model represents the real soil behavior and the static pile capacity is calculated from it. Pile discontinuities like necking, bulbs, and material changes introduce stress wave reflections, which can influence the calculated signals strongly. Reflections from bulbs yield an almost similar wave equation result as a local stiff soil layer and a necking similarly results as a local soft layer. When pile discontinuities are not properly taken into account, either a proper match cannot be obtained or the capacity prediction will not be reliable. Soil properties can be confused with pile discontinuities.
For DLT on cast in situ piles a drop hammer with a guiding system has to be mobilized. The required ram mass is as rule of thumb 2% of the maximum load that has to be applied. A crane is required to move the drop hammer over the building site. The pile head has to be prepared to prevent damage from impact loading. An epoxy or grout cement is used to smooth the pile head surface to prevent stress concentrations during impact loading. The location of the transducers has to be at least 2 pile diameters from the pile head. When the pile head is located at ground level this requires an extension of the pile head for a similar length or the excavation of the pile head. For small capacity piles multiple piles can be tested in one day. For loads above lOMN the testing rate is normally in the range of two piles per day. For STN a loading device with a reaction mass catching system has to be mobilized. The required reaction mass is as rule of thumb 5% of the maximum load that has to be applied. Local available material can be used as reaction mass to reduce transport costs. A crane or a crawler system is required to move the STN device over the building site. For loads up to 4MN a STN device with a hydraulic catch mechanism can be applied. For higher loads STN requires a gravel catch system. Testing can take between 0.5 and 2 days per pile depending on the pile capacity. However for piles with a capacity less than 4 MN, a loading device with hydraulic catch mechanism can be applied and the number of piles tested in one day are in the same range as with DLT. STN can be even more efficient when the loading device and hydraulic catch mechanism are placed on crawlers. An epoxy or grout cement is used to smoothen the pile head surface to prevent stress concentrations during push loading. 2.6 Chance of pile damage
2.4 Reliability for testing on cast in situ piles With DLT the load on the pile head is introduced by an impacting ram. When the ram is not properly guided and hits the pile in an eccentric way, bending stresses will occur and result in excessive compression and/or tension stresses which can damage the pile. Most cast in situ piles need considerably more displacement to mobilize the ultimate capacity than driven piles. This softer response will easily gener-
Because of the many unknowns that have to be solved to perform a proper DLT signal matching analysis on cast in situ piles, there is considerable chance of errors in pile capacity predictions. The load measurement for STN is similar as for static load testing and unknown pile properties of cast in situ piles will not influence the load meas629
For STN the load displacement behavior is calculated in most cases by the Unloading Point Method, however signal matching techniques are also applied. 3.1 Accuracy With DLT on pre-cast driven piles, the load in the pile is measured by strain transducers mounted on the pile shaft. Pre-cast piles are considered to be of homogenous material and with the method described in section 2.2 and based on the determination of the stress wave the E-modulus can be determined accurately. The toe reflection will be visible at several stages of driving and the stress wave velocity can be determined easily. Only when the pile head is heavily reinforced will the E-modulus at the pile head be different from the E-modulus calculated by the stress wave velocity. With STN the load is accurately measured by a calibrated load cell placed on the pile head. The measured load is not dependent on the pile properties. The load measurement error is less than 0.1% of the maximum range of the load cell. Fig 9. Dynamic load test on a pre-cast pile 3.2 Reliability ate tension waves. Cast in situ piles are not designed to withstand high tension stresses. As soon as allowable tension stress levels are reached the impact energy has to be reduced to prevent pile damage. As a result, DLT has to be stopped at a stage where full capacity has not yet been mobilized. With STN the duration of the loading is long enough to keep the pile is under constant compression and tension stresses will not occur. To prevent bending stresses the piston of the statnamic device is installed exactly on or near the center of the pile head cross section. The launching of the reaction mass, and the resulting push load starts from the center of the pile. 3 THE APPLICATION OF DLT AND STN ON PRECAST DRIVEN PILES For pre-cast driven piles both DLT and STN are performed after a setup period after pile installation. This allows the soil to recover from driving induced disturbances like pore water pressure. In most cases the soil will regain strength during the setup period. For DLT strain and acceleration transducers are mounted on the pile shaft near the pile head. The load displacement behavior is calculated by signal matching. 630
The capacity of driven piles is mobilized at relative small displacements. Both DLT and STN are performed after a set up period. For DLT the pile load displacement behavior is calculated by a signal matching technique (CAPWAPTM,TNOWAVE) in most cases. For STN the pile load displacement behavior is determined by a direct method, the Unloading Point Method (UPM) and in some cases by signal matching. 3.3 Economy DLT has the advantage that the pile driving hammer used for pile installation can also be used for redriving the piles after a set-up period. However, when the pile driving hammer has to be used for constant production, an additional pile driving hammer or drop hammer has to be mobilized. When the mobilization of the full pile capacity is requested, the production hammer might not be sufficient to mobilize pile capacity after the set up period and an additional heavier hammer has to be mobilized. For STN the same economical conditions are applicable as mentioned in paragraph 2.5 (Ginneken, van G.J.J., 2000)
Table 1. Preferences for DLT or STN with respect to economy for driven piles Driven piles Preferred Soil set up I DLT I STN dium medium to high
**
*****
STN
4 CONCLUSIONS For bored concrete piles, auger piles and caissons the dynamic load testing method has some disadvantages and is less suitable and statnamic load testing is the preferred method. The most important reasons for the preference of statnamic load testing in the case of cast in situ piles are:
1. Accuracy in load measurement STN is not dependent on pile material and cross section properties 2. No influence of cross sectional variations STN results are not influenced by cross sectional variations over the pile length 3. No tension during compressive testing STN long duration loading will keep pile under constant compressive pressure 4. Concentric loading Easy placement of STN loading device in center of the pile 5. Pile and soil response closer to static With STN the pile moves as one unit, similar to static load tests. Stress wave phenomena can be neglected resulting in a simple method of analysis
3.4 Mobilization of capacity Set up phenomena can increase the soil resistance considerably. The pile driving hammer used for pile installation might not be able to mobilize the full pile capacity in such a case. Another reason that capacity can not be mobilized with DLT is that the load cannot be increased because compression or tension stresses becoming too high. To mobilize the pile capacity a STN device will be sent to the building site with at least a corresponding loading capacity. Only when the piles are over-designed will the full bearing capacity not be mobilized.
For driven piles both DLT and STN methods can be applied reliably and each has its advantages and disadvantages. A big economic advantage for DLT can be the use of the production rig for testing. A big advantage for STN is the fact that maximum available energy can be used to mobilize capacity and that that testing does not have to be stopped when tensional stresses become too high like with DLT.
3.5 Chance of pile damage For DLT there are some cases with a chance of pile damage. In the case of low friction and a soft toe response tension waves will be generated during DLT. When the maximum allowable tension stresses are reached the load on the pile cannot be increased because this will generate higher tension stresses and the pile will experience damage. In the case of a pile with a hard toe response, for example pile toe on rock, the compressive stresses at the pile toe can theoretically be two times higher than the maximum compression stress at the pile head. This is caused by the superposition of compression stress waves at the pile toe. So, if during DLT the compression stresses at the pile head are higher than half the compressive strength of the pile material, collapse of the pile material at the pile toe will occur. In this case piles can only be tested up to half the compressive strength of the pile material, which may not correspond with the capacity of the pile.
5 REFERENCES Abbasi, A.F., Al-Tayyib, 1990. Effect of hot weather on pulse velocity and modulus of elasticity of concrete. Materials and Structures, 1990, 23, ~~334-340 Ginneken, van G.J.J., 2000, Introduction Statnamic Load Testing in Europe, Case Studies in the Netherlands. Proceedings of the Sixth International Conference on the Application of Stress-wave Theory to Piles, Sao Paulo. Holeyman, A.E., 1992. Keynote Lecture: Technology of Pile Dynamic Testing. Proceedings of the Fourth International Conference on the Application of Stress-wave Theory to Piles, The Hague, F.B.J. Barends, Editor, A.A. Balkema Publishers, ppl95215.
For STN the pile is kept under constant compression and tension waves are suppressed. Superposition of compression waves at the pile will not occur. As with SLT piles can be tested near to the compressive strength of the shaft.
Franklin, R.E., King, T.M.J. 1971, Relations between compressive and indirect-tensile strength of
631
concrete, Road Research Laboratory, RRL Report LR 412 Madan B. Karkee, Takashi Horiguchi, Hideaki Kishida. Static and Dynamic Tests for Evaluation of the Vertical Load Bearing Capacity of Piles. , 22nd DFI Annual Member's Conference, Toronto, Canada, 1997, pp199-214 Sparks, P.R., Menzies, J.B., 1973. The effect of rate of loading upon the static and fatique strength of plain concrete in compression. Magazine of Concrete Research, Vol25/ 1973, No. 83, pp 73-80.
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10 Case histories, pile set-up and correlations between test methods Prediction reliability
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Application of Stress-Wave Theory to Piles, Niyama & Beim (eds) 02000 Balkema, Rotterdam, ISBN 90 5809 150 3
Keynote lecture: Improving the reliability of pile bearing capacity prediction by the dynamic increasing energy test (DIET) Nelson Aoki University of SLio Paulo, SLio Curlos, Brazil
ABSTRACT: The dynamic increasing energy test (DIET) allows the measurement of the dynamic resistance displacement curve and the prediction of the static resistance - displacement curve of the pile-soil system. The variations of energy and work done are evaluated by Hamilton’s principle of energy conservation, and this evaluation allows the prediction of the dynamic failure load of the system. The usefulness of this approach is confirmed by the good correlation between the results of the DIET and of the static load test. static approaches and the definition of the ultimate resistance, are of foremost importance. This paper will deals with: (a) the ultimate resistance definition in the conventional static load test and in the constant energy dynamic test, accounting for the energy in the system; (b) the Hamilton’s principle of energy conservation (Clough & Penzien 1975); (c) the analysis of the dynamic resistance displacement curve of the dynamic increasing energy test and (d) some case histories. The static pile bearing capacity, in the traditional dynamic loading test, is predicted from the analysis of one blow, among several blows of same energy. This approach can be called a constant energy test. In that case, depending on the magnitude of the observed permanent displacement (set), the Smith’s model static resistance (Smith 1960) could satisfy either the lower or the upper bound theorems of the limit analysis (Chen & Liu 1990). The dynamic increasing energy test (DIET) involves the measurement of the total resistance to penetration or simply dynamic resistance, the displacement and the kinetic energy applied to the system at increasing energy hammer impacts (Aoki 1989b). The result is a dynamic resistance - displacement curve, from which the corresponding static resistance - displacement curve (Smith’s model) is generated. The dynamic resistance - displacement curve is then examined and the variations of energy and work done are analyzed by using the Hamilton‘s principle of energy conservation. Under the action of the failure load the system displaces as a rigid body. Considering the similarity of the failure under static and dynamic loading conditions, since both events involve energy and work done variations, it is possible to evaluate the ultimate
1 INTRODUCTION The main purpose of a dynamic pile test is to predict the static response and it is essential that any transfer model can be applied to both static and dynamic loading conditions (Randolph & Deeks 1992). This is fimdamental for reliable predictions of behavior under static and dynamic conditions (Aoki 1989a). The evaluation of the static resistance by the application of the stress wave theory to one blow of hammer is questionable and some discussion points are: (a) the choice of one blow among several impacts of nearly the same energy; (b) the effects of the elapsed time, from the end of driving to the start of the test; (c) the magnitude of deformations around the pile toe and shaft; (d) the comparison of the dynamic penetration with the static movement; (e) the methodology of the static load test and the evaluation of the pile bearing capacity (Fujita & Kusakabe 1988). The need for the standardization of the pile - soil interaction model, the establishment of soil parameters, the statistical treatment of discrepancies between the model and reality and derivation of design values for the load bearing capacity, which can be used within the modern Ioad and resistance partial factors design approach, are essential for the total foundation quality assurance (Vrouwenvelder & Dieterman 1992). The average resistance and the standard deviation allow verifying the actual characteristic load of the piling and the quality assurance depends on how and who takes the decision to stop the driving in the field. In contrast, the safety in the allowable load philosophy is simply a matter of personal past experience. In any case, the reliability of the pile bearing capacity prediction, by dynamic or 635
bearing capacity with the dynamic increasing energy test (Aoki 1997). The DIET has been largely used in Brazil as an alternative to the traditional static load test and the constant energy dynamic test (Aoki & Mello 1992, Beim & Aoki 1996). The static and dynamic results compare reasonably well (Aoki 1989c, Niyama & Aoki 1991). The DIET can also be used to analyze time effects on pile behavior “set up” after short time intervals. 2 RESISTANCE - DISPLACEMENT CURVE AND THE ULTIMATE PILE RESISTANCE The goal of a loading test is the determination of the ultimate resistance or bearing capacity of an isolated foundation pile. The isolated foundation pile is a system constituted by the structural pile element and the soil layers along the shaft and under the pile toe. This system’s response to the action of an axial load applied to the pile head is the displacement of this point, which is a complex function of many independent variables. Among them, the most important are: (a) time history of the applied force; (b) geometry of the pile and the soil layers; (c) physical properties of the pile and the soils; (d) peculiarities of pile installation that change material’s properties; and (e) elapsed time after the pile installation. During the static loading the applied force and the vertical movement are measured at each loading stage. In this case the applied force is equal to the mobilized resistance and the response of the pile soil system to the static loading is the well-known load - settlement curve. The slow maintained loading (SML), the quick maintained loading (QML), the constant rate of penetration (CRP) and the cyclic loading are examples of static test procedures (Fellenius 1980, Whitaker & Cooke 1961). Nowadays, the definition of the static pile bearing capacity or static ultimate resistance, which is the maximum reaction capacity of the pile - soil system, is a matter of discussion in the pile foundation engineering practice. Nevertheless, it is recognized that the shape of the load - settlement curve should be included in this definition (Reese 1972, Fellenius 1980) and this can be done by the consideration of the deformation energy of the pile - soil system. For a given curve there are many interpretations, each one leading to a particular ultimate resistance. The absolute or relative displacement values and the relationship between the elastic and permanent components of the displacement after the unloading or the time dependent behavior are some of the criteria for the establishment of the static bearing capacity (Vesic’ 1975, de Beer 1988). The absolute displacement equal to 10% of the pile dimension is a typical definition of the ultimate resistance (Terzaghi 1942).
636
The upper bound and lower bound theorems are used to determine the static failure load from energy transformations and equilibrium of forces considerations in the theoretical soil mechanics limit analysis (Chen & Liu 1990). By the upper bound theorem, the load is not less than that of the actual failure load, determined by equating the external rate of work to the internal rate of energy dissipation in an assumed velocity field which satisfies: (a) the velocity boundary conditions and (b) the strain and velocity compatibility conditions. Note that the stress distribution in the system does not need to be in equilibrium. Under this limit load the system loses its capability to store more elastic recoverable energy of deformation and starts to displace as a rigid body. By the lower bound theorem, the stress field must satisfy the stress boundary conditions and must nowhere violate the materials yield condition. The most popular dynamic loading test is the impact of a hammer on the cushion - pile - soil system. In this case the pile bearing capacity can be predicted by the old pile dynamic formulae or by the application of the stress wave theory to the hammer impact (Whitaker & Bullen 1981, Goble et al. 1975). The statnamic test has been developed as an alternative dynamic test (Middendorp 1993). During the hammer impact the time histories of deformation and acceleration are recorded at the instrumented section and the velocities and displacements are calculated by integrating the measured accelerations. The force history is obtained from the deformation and the applied kinetic energy is evaluated by integrating the force and the displacement. The dynamic resistance or total resistance to the pile penetration is evaluated from the analysis of forces and velocities. The correspondent static resistance is evaluated by using the classical Smith‘s model (Smith 1960) or the derived model of CASE and CAPWAP methods (Goble et al. 1981, Rausche et al. 1985, Goble & Likins 1996). The ultimate resistance from one single hammer impact is defined by the Davisson‘s limit load criterion (Davisson 1972). By this criterion the ultimate resistance is defined by the elastic toe displacement equal to (0.15 + B/120) inches, where B is the toe dimension. Finally it can not be forgotten that the soil resistance changes with the time and neither the static nor the dynamic loading can predict the future behavior of the pile - soil system, from the measured values at the time of the test. It can be concluded that the definition of the ultimate resistance is not clear in both static and dynamic loading tests. 2.1 Static force - displacement curve response. The loading is called static when the force (Q) is reached in stages of infinitesimal increments (dQ) and each stage is of infinite time duration. Thus, the
frequency of static loading is equal to zero. However, in practice, there is no static loading once the increment of force (AQ) and the time duration (At) have finite values. Figure 1 presents the static force - displacement curve, resulting from the force increment (AQ) and displacement increment (Ap) and the consequent increment of the deformation energy (AV).
if the zero load condition is reached in the inverse path described for the static loading. In the actual example of a building construction, the unloading path does not exist while the load-unloading cycle is observed only for the live load. The potential energy of deformation (V) stored in the system after the static loading can be elastic recoverable or permanent, depending on the rheological characteristics of the materials. This fact can be only known after the system unloading, when the potential energy of deformation (V) is transformed into recovered elastic energy (V,) plus the work done in the system (W,,) in such a way that
v = v, w,,
(5)
f
Figure 1. Static force - displacement curve.
where V = stored potential energy of deformation, V, = recovered elastic energy, and W,, = work done by the non-conservative forces. The reliability of the ultimate resistance predicted from the static load - settlement or the resistance displacement curves is discussed thereafter.
The rate of internal energy of deformation corresponding to the force increment (AQ) is the power
2.2 Shape of the static load - settlement curve.
g=AV I At (1) where 9= power or rate of the internal energy of deformation increment, AV = increment of energy deformation, and At = increment of time loading. The average power of the actual static or dynamic loading is variable from some watts to various kW. Under the action of the static load (Q), the displacement of the pile head (point D) is expressed by
P = Pe + Pp (2) where p = pile head displacement, pe = elastic displacement component, and pp = permanent displacement component. The nature of the deformations can not be known at the end of the loading stage. The potential energy of deformation (V) at the end of the loading is the hatched area ODFO. The complementary energy (Langhaar 1962) can be calculated by the difference
0
Vc= Q . P - V (3) where V, = complementary energy (area OGDO), Q = applied force, p = total displacement of the pile head, and V= potential energy of deformation. The static equilibrium of forces is expressed by R=Q
Figure 2 presents the typical static load -settlement curve in the case where the ultimate resistance is well defined by the final shape of the curve. This curve correlates the settlement (p) to the load (Q) where it is implicit that the static reaction (R) is always equal to the static load (Q). The failure is characterized when the pile head displaces from point D to point A under the action of the constant limit load (Q,), equal to the ultimate system resistance (Ru). QU
Figure 2. Load - Settlement curve: defined ultimate resistance.
When the limit load (Q = Q,) corresponding to the point D of the loading curve is applied, the limit total displacement is (p = pu), the limit potential energy of deformation is (V = V,), and the limit complementary energy is equal to (V, = V,,). In the case where all materials are perfectly rigidplastic, the unloading path is the horizontal line DF. In that case the deformation energy is completely transformed into work done by the resistant forces
(4)
where R = total mobilized resistant forces (resistance or reaction), and Q = applied force (load or action). According to expression (4), the system response under static loading can be indifferently expressed by the relationship between the force - displacement curve or the resistance - displacement curve. The system response to unloading is called static 637
(V = Wn,). There is no elastic recovery (Pe = 0) and the total settlement OB is permanent (p = pp). For perfectly elastic materials the loading path is the curve OD and the unloading path is DO resulting in a minimum work done (Wnc= 0) and (V = Ve). In actual cases some elastic and plastic deformations take place and the unloading path is DE. The recovered elastic energy (Ve) is represented by the area DFED and the area ODE0 is the work done (W,,) in the system. For any point on the vertical asymptote line between points D and A, all the additional energy of deformation is fully transformed into work done, if the elastic recovery energy becomes constant after point D (DE line being parallel to AC line). It is not possible to know beforehand if this defined failure load is equal to the upper or to the lower bound of the failure load, unless more data concerning the external rate of work and the internal rate of energy dissipation could be known. Figure 3 presents the typical shape of loadsettlement curve corresponding to the case where the resistance reaches a maximum equal to Q m a x at point F and thereafter it drops up to a constant residual ultimate resistance, after point D. The limit load is again the well defined value Qu on the vertical asymptote DA to the load-settlement curve.
where Q = applied force, Q, = physical ultimate resistance, and p = settlement. Expression (6) can also be written lim Q = Q ,
(7)
V+m
where V potential energy of deformation. In this case the load - settlement curve can be described by the well-known expression (Veen 1953) =I
where Q = applied force, Qu= physical ultimate resistance, a = curve shape coefficient, and p = settlement. This expression can also be rewritten as
(9)
The potential energy of deformation is equal to (10) J
L
where V = stored potential energy of deformation, Qu= physical ultimate resistance, p = settlement, and a = curve shape coefficient.
0
Q"
0
--P
load
+-'
)
-s5 2
C
ID
-----
B
D
I
Figure 3. Load - settlement curve: defined residual ultimate resistance.
Figure 4. Load - settlement curve: physical ultimate resistance.
This shape of curve is observed in the cases of floating piles where, after the peak of resistance (Qmax) the system resistance drops up to the defined residual ultimate resistance (QL,).The unloading path is dependent on the pile-soil system characteristics as explained in the case of figure 2. Figure 4 presents the case where the ultimate resistance is not defined. The system displacement increases with increasing loading and the vertical asymptote to the curve appears to be at the infinity. In this particular case the ultimate resistance can be taken as the physical ultimate resistance defined by the condition
Considering expression (3), the complementary energy can be estimated by the expression
where V, = complementary energy of deformation, Q, = physical ultimate resistance, p = settlement, and CI = curve shape coefficient. The limit complementary energy corresponding to the physical ultimate resistance would be lim V,= V,,
(12)
P+W
lim Q = Q,
where V,,
P+a
638
= limit
complementary energy.
When the curve is expressed by Veen's expression (8), the limit complementary energy becomes constant and its value is expressed by
v,,
= -Q
the structural pile material strength. This is expressed by the following condition: (14)
QSQu 200
5.6
3 STATIC LOAD TESTS Two slow maintained load (SML) tests were carried out in Aug-Sep/98 according the Brazilian code NBR-12131. The penetrations were 6.0 and 7.0 m for the CFA-1 pile and the CFA-2 pile respectively. The measured load-displacement curves are presented in Figure 2. The tests procedures, instrumentation setups and interpretation are described by Kormann et al. (1999). The following points were outlined by these authors: 708
Figure 2. Static load tests results.
5 . The performing of additional SPT borings just beside the piles. These investigations and the results of all the available data are presented and discussed in the following items. 4 LOW STRAIN TESTS Since the penetration of the piles was well controlled, the low strain tests allowed to calculate the wave speed, resulting 3550 m/s and 3800 m / s for the CFA-1 and the CFA-2 piles respectively. These wave velocities are somewhat reduced. Although the concrete used in both piles was essentially the same, the higher wave speed of the CFA-2 pile might be associated to a material having a slightly better performance (Kormann et al., 1999).
5 HIGH STRAIN DYNAMIC TESTING In order to perform the dynamic testing the pile tops needed to be reinforced. The reinforcement was carried out using longitudinal steel bars and a cylindrical steel ring mounted at the top of the piles (Fig. la). It is important to note that the reinforcement reduced the pile embedded length by approximately one meter which resulted in some loss of shaft friction near the top of the piles. The dynamic load tests were conducted in Dez/99. For a proper selection of the hammer to be used for dynamic testing of cast-in-place piles, Hussein et al. (1996) suggest a ram weight about 1.5% of the pile static resistance to be verified. A 30 kN free fall hammer was selected for the high strain dynamic tests. One cushion made by wood plate was used to reduce the impact maximum stress. Two piezoeletric accelerometers and two strain transducers were attached to the piles. A Pile Driving Analyzer, PAL model, was employed (PDI, 1999). In order to mobilize as much as possible the soil resistance, the tests employed increasing hammer energies (e.g. Aoki 1989, Bernardes 1989, Beim & Aoki 1996). The drop height was increased in 20 cm increments until the compressive stresses
caused failure of the pile material. The set and rebound of the pile top were recorded for all blows. Table 2 summarizes the dynamic load test data for the CFA-1 and the CFA-2 piles. The first pile tested was the CFA-2 pile. At the beginning of the test the new cushion with low stiffness and low coefficient of restitution generated a low driving system efficiency - a reduced energy (EMX) was transferred. The pile was initially stroked six times with 40 cm fall height until the set became perceptible. During the test, the driving system efficiency increased with the succession of the blows. An efficiency ranging from 12.2 to 38.6% was observed. The dynamic load test comprised seven blows with drop heights ranging from 60 to 170 cm. The maximum tension stresses below sensors (TSX) were small for all blows. The maximum compressive stress (CSX) of 28.5 MPa generated by the seventh blow caused the pile top material to fail just below the steel reinforcement. The CFA-1 pile received six blows with drop heights comprised between 40 and 140 cm. The driving system efficiency increased from 1 1.7 to 32.1% during the test. The tension stresses were slightly higher than i n the CFA-2 pile, but as expected the successive increase in compressive stresses caused the failure of the pile top immediately below the steel ring. The compressive stresses that controlled the end of the test (15.0 MPa at the gages) were lower than the CFA-2 ones. It can be related to bending effects due to eccentric blows or concrete properties. Figure 3 shows the records of force and velocity multiplied by the impedance for the two piles. For the CFA-1 pile, a wave speed of 3000 m/s is suitable. The CFA-2 pile shows a wave speed of 3250 d s . These findings are in full agreement with the wave speeds measured in the PIT tests, which due to the level of deformation usually present higher values. The ratio between the wave speed from both tests for the CFA-2 pile to the CFA-1 one was the same regardless the type of test (1.07- 1.OS). Due to the reduced length of the piles, the width of the pulse impact is considerably higher than 2L/c. Despite that, it will be demonstrated that good results can be achieved. The traces of the force and the velocity of the CFA-1 pile suggest a high skin friction (Fig. 3a). The succession of the increasing energy blows causes the force signal to be close to the velocity signal after 4L/c. It can be related to increased pile point displacements. A considerable pile set (9,4 mm) was measured at the last blow (Table 2). It is expected a high mobilization of the soil resistance. The first record of the CFA-2 pile (Fig. 3b) shows an increase in the velocity just after 2L/c. It is related to a hammer interference. The subsequent records do not present such feature.
709
Table 2. Dynamic CFA- 1 Blow Drop number height (cm) 1 40 2 60 3 80 4 100 5 120 6 140
load test data.
EMX CSX (kNm) (MPa)
TSX (MPa)
Set (mm)
Rebound (mm)
1.4 3.4 5.1 8.2 11.8 13.5
0.4 0.4 0.3 0.5 0.1 0.7
0.1 1.8 3.0 5.2 8.0 9.4
1.9 2.2 2.2 2.3 2.0 2.1
9.4 11.2 12.5 13.8 15.1 15.0
CFA-2 Blow Drop number height (cin) 1 60 2 80 3 100 4 120 5 140 6 160
EMX CSX TSX (kNm) (MPa) (MPa)
Set (mm)
Rebound (mm)
2.2 5.0 8.2 12.2 14.9 18.1
0.5 0.6 1.4 2.2 2.8 4.3
1.6 2.9 3.8 5.1 4.2 3.7
14.1 19.0 21.5 23.9 25.1 27.4
0.0 0.0 0.0 0.5 0.3 0.0
BN - Blow number H - Hammer drop height
0 - Pile ( ~ u c ) 1500 00
BN=l
2.70
H=40cm
4 1s
-7
-7 3 40 rns
+]
1500.00
L
_.--
B N = 2 H=60cm
I
/
\ .-
4
L /
3 20 rns 3 20 rns
I
I
'1
-\ I
'
,
\ .//
A
2
0 5m5 - 20 5m5
Figure 3. Records of force and velocity multiplied by the impedance for CFA-1 (a) and CFA-2 (b) piles.
710
The force signal is considerably higher than the velocity record. It can be due to compressive upward waves that are related to a high skin friction. The comparison between the measured sets and rebounds of the two piles also reflects the distinct behaviors observed both in the static load tests and in the dynamic records. Despite the higher energies applied to the CFA-2 pile, its sets are lower than a half of the CFA-1 ones. The measured rebounds of the CFA-2 pile were as high as 5.5 mm, while in the CFA- 1 pile the maximum rebounds were 2.3 mm. All blows presented in Figure 3 were submitted to CAPWAP analysis. The results will be presented in the item 9.
6 VISUAL SHAFT INSPECTION After the dynamic tests, the soil around the piles was excavated approximately 3.2 m below the ground level. The visual inspection of the soil allows the following remarks to be made: 1. As expected, the silty clay around the piles has polished and shiny surfaces closely spaced. Slickensides following a preferential orientation also were found. The clayey soil around the two piles seems to be similar. 2. A hard stratum of sandy silt is present from the ground level to 1.3 m deep at the CFA-2 pile (see Fig. la). It is important to note that skin friction of this layer acted only on the static test, since the soil around the pile top was removed for the dynamic testing. The following comments can be outlined about the inspection of the piles geometry and material: 1. The surface of the two piles was very smooth and regular. There was no sudden changes in the pile diameter which was expected for this type of pile construction in stiff clayey soils. 2. The actual diameter of the two piles is lower than the nominal diameter (35 cm). The measured diameter ranged between 33 and 34 cm. 3. The reports of the installation data of the two piles indicated excessive amount of concrete used 47-59% more volume than the nominal volumes of the piles. The direct measurements of the diameter of the piles show that there is no reason for the high concrete overconsumption and a better control during pile installation is required. 4. The CFA-1 pile shows some small soil intrusions along the shaft. They are present at the external pile surface and also approximately over to 4 cm inside the concrete mass. The CFA-2 pile does not exhibit soil intrusions. It may explain the lower wave speed (3000 d s ) measured in the CFA-1 pile.
711
7 CONCRETE TESTING In order to evaluate the concrete density, a total of 17 samples of the material of the two piles were collected just below the steel reinforcement and at the bottom of the excavation. Table 3 presents the mean results of the tests. It is important to note that the concrete specific wei ht of the piles - in the range of 19.13-19.52 kN/m - is low. It is due to the high porosity of the material. These findings are in agreement with the low wave speeds measured. It can be noted that despite the low concrete density and probably the low resistance the piles supported well the loads both of the static and the dynamic tests.
5
Table 3. Average pile concrete properties. Pile Concrete specific Voluine of voids / Toweiglit (W/m’) tal volume (9%) CFA- 1 19.52 25.07 CFA-2 19.13 25.56
It should be noted that the concrete specific weight is currently used in wave equation analysis to calculate the dynamic elastic modulus of the pile. It is a common practice to adopt to the specific weight the value of approximately 24 kN/m3. However, if the cast-in-place shaft exhibits a reduced specific weight the analysis would produce overestimated results. In order to illustrate the influence of the concrete specific weight, Figure 4 shows a simplified analysis for the CFA-2 pile. The pile capacity was evaluated with the Case Method. It is observed that for a specific weight range of 20-24 kN/m3 the pile capacity can change about 20%.
Figure 4. The influence of the concrete specific weight in the pile capacity evaluation.
8 ADDITTIONAL SPT BORING Since the excavations around the piles were not sufficient to explain the distinct behavior between the two piles, complementary SPT borings were performed just beside the CFA-1 and the CFA-2 piles. The distance between the borings and the piles was approximately 50 cm. Obviously, the excavation around the piles implies in a stress relief that
Table 4. Complementary SPT boring data. CFA- 1 CFA-2 Depth Blows130 cni Blows/30 cm Blows130 cm Blows130 cm below (first and (second and (first and (second and third 15 cm) second 15 third 15 cm) ground second 15 level cm) NSI’T cm) NSPT
(m) 4.0 5.0 6.0 7.0
16 21 28 30129cni
27 30128cm 30123cm
*
16 22 19 17
27 30 1 2 5 cni 30 19
changes the conditions of the soil. In effect, even the pile installation changes the natural state of stresses. However, since the conditions of the ground around the two piles were similar at the time of the complementary SPT boring, a qualitative comparison can be made among them. Table 4 presents the results. The following remarks can be made: 1. The complementary SPT borings data are in agreement with the SPTT-1 information. 2. The new borings revealed a lens of silty sand just above the point of the CFA-2 pile and 60 cm below the CFA-1 pile. These thin layer is approximately 40 cm thick and it probably appears in the soil profile as presented in Figure 1 a. 3. The N s p ~values along the shaft of the CFA-2 pile are not higher than the CFA-1 ones. At the point of the CFA-2 pile the N s p ~values are lower than the CFA-1 ones. 4. On the basis of the known limitations of the SPT testing, the blow count suggests that the higher resistance of the CFA-2 pile could not be explained only by a stronger soil around the shaft or at the pile point. The hypothesis of the influence of the grout pressure in the performance of the pile cannot be disregarded. Also, since the sand lens intercepts the CFA-2 pile just above the tip, a distinct response of this soil to the installation procedure - such as an slight enlargement - might affect the pile end bearing. 9 CAPWAP ANALISYS: COMPARISONS WITH STATIC LOAD TESTS The records of both the CFA-1 and CFA-2 piles were submitted to conventional CAPWAP analysis. In order to perform a suitable analysis eight pile and soil segments were employed. A concrete specific weight of 20.0 kN/m3 was adopted. The main results are presented in Table 5. The skin quake values of the two piles lies in a narrow range (0.97- 1.622 mm) and they do not depend on the energy level. On the other hand, the toe quakes exhibit a trend to increase with the energy transferred to the pile. It can be related to a growing mobilization of the toe resistance. High toe quakes
were found, in special for the CFA-1 pile. For the two piles the ratio between the toe quake and the measured set is approximately constant - a representative range is 1.O- 1.7. The Case Method damping JC values of Table 5 does not exhibit a trend and should be disregarded. The shaft and toe Smith damping values of the two piles showed dependent on the energy level. At the first blows the toe damping was too high. It decreases with the increasing of the drop height. Such behavior is an agreement with the findings of Aoki & de Mello (1992). These authors show that the damping and quake of the Smith model are not constant soil parameters, but are dependent on the energy level. The use of the radiation damping - that may account for the soil motion associated with the pile movcment in drilled shafts - did not lead to improved results. It seems in agreement with the smooth surface of the piles revealed by the visual inspection and the stiff characteristics of the soil. Rausche et al. (1996) discusses the applicability of the Multiple Blow Analysis (MBA). They recommend such analysis for variable energy blows and provide an example for a drilled shaft. However, for the piles here studied such analysis did not seem suitable. The MBA analysis keeps constant the end bearing. That is not in agreement with a expected increased mobilization of the toe resistance due increasing hammer energies. In this study, the Residual Stress Analysis (RSA) option also did not improve the results. The mobilized capacities of the two piles are compatible with the results of the static load tests. At the end of the dynamic tests, for the CFA-1 and the CFA-2 piles respectively, the maximum mobilized capacities were 10.7% and 48.1% higher than the static capacities reported by Korrnann et al. (1999). The high CAPWAP resistance of the CFA-2 pile suggests that the limited settlements of the CFA-2 pile in the static tests may have conducted to an underestimation of the extrapolated capacity. It is important to note that the CFA-1 and the CFA-2 piles were statically loaded before the dynamic test. These preloading leads to a set of residual stresses both in the pile and in the soil (e.g Massad 1992, Maiorano et al. 1996). The comparison between static and dynamic load tests performed on the same pile should account for the residual stresses. The set of residual stresses does not change the ultimate load. However, the form of static loading curves can be markedly affected. The residual loads generated by a cycle of static or dynamic loading increase the pile head stiffness. In addition, as discussed by Dkcourt (1998), the meaning of the physical failure cannot be applied to bored piles. Cast-in-place piles may develop increasing point resistance over to large displacements.
712
Table 5 . Main CAPWAP results. CFA- 1 1 Blow number Mean Shaft Quake (mm) 1.298 0.764 Toe Quake (mm) 1.298 Shaft Damping (s/m) 1.419 Toe Damping (dm) 0.36 Jc 1.02 MQno 62 1 RS (kN) 24 8 RT (kN) RU (kN) 869 RS - Skin mobilized resistance
CFA-2 6 1 2 4 3 2 3 4 5 0.996 1.001 1.001 1.001 1.000 1.361 0.9 7 1.445 1.526 2.682 4.080 6.413 8.557 11.05 0.757 1.004 2.109 3.422 1.183 0.713 0.533 0.39 0.512 0.324 0.208 0.091 0.099 4.936 2.526 2.055 1.301 2.682 1.411 0.677 0.663 0.13 0.68 0.7 1 0.42 0.41 0.07 0.00 0.18 1.oo 1 .oo 1.28 0.86 0.93 1.32 1.37 2.34 1.96 2.21 1.58 800 1296 1592 1570 1470 590 620 719 760 314 148 167 280 300 300 180 296 310 886 900 1019 1060 1114 1444 1759 1750 1780 RT - Toe mobilized resistance RU- Total mobilized resistance
If the soil resistance degradation effects are reduced, it is expected that the succession of static or dynamic cycles of loading will lead to increased pile capacities. For instance, Liu et al. (1996) points out that when the dynamic test is performed after the static loading, in average a 10% increased pile capacity can be found. On the basis of the discussion above, it seems suitable to compare static and dynamic tests performed at the same pile as a succession of cycles of loading. In order to perform such analysis, the permanent displacement at the end of a cycle of loading can be added to the next one (e.g. Seitz 1984, Niyama & Aoki 1991). Figure 5 shows the static load tests and the CAPWAP static simulations superimposed as cyclic loading. In order to plot the calculated curves, the measured sets were employed. The range of 0.928I . 1 15 is representative for the ratio among the measured sets and the CAPWAP ones. The unloading of the blows number 4 to 7 of the CFA-2 pile was adjusted as a straight line between the CAPWAP maximum displacement (DMX) and the measured set. The simulations plotted in Figure 5 shows a remarkable agreement with the static load tests. The CAPWAP inferred static behavior follows within reasonable limits the trend of the static load tests. For the CFA-I pile simulations, it can be observed that the succession of blows increases the end bearing mobilization and reduces the toe stiffness. The gradient of the last straight portion of the loading cycles increases with the succession of the blows. In effect, the ratio RT / Toe Quake decreases when the drop height increases. It can be related to a nonlinear stress-strain behavior of the soil. Despite the lower displacements, the same trend can be observed in the CFA-2 pile data.
10 CONCLUSIONS At the current stage of the research here presented, the following conclusions can be outlined: 1. The visual shaft inspection of the two CFA piles revealed a very smooth and regular surface. It is consistent with the stiff and hard clayey soil 713
5 1.595 3.573 0.138 1.371 0.84 1.14 1574 300 1874
6 1.622 5.866 0.1 17 0.974 0.63 3.26 1776 309 2085
7 1.000 5.028 0.1 17 0.702 0.59 3.12 1876 306 2182
around the piles. The actual diameter of the two piles, ranging between 33 and 34 cm, is lower than the nominal diameter (35 a n ) . It shows that there is no reason for the high concrete overconsumption reported by the installation monitoring system. The CFA-I pile shows some small soil intrusions over to 4 cm inside the concrete mass. The wave speed of the CFA-1 pile (3000 m/s) was lower than the wave speed of the CFA-2 pile (3250 m/s). 2. The concrete testing of the CFA-1 and the CFA-2 piles showed a reduced specific weight (19.52-19.13 kN/m3) and a high porosity. Despite that, the structural behavior of the piles was satisfactory both in the static and dynamic load tests. However, it should be noted that the concrete specific weight is currently employed in wave equation analysis to calculate the dynamic elastic modulus of the pile. The use of an overestimated concrete specific weight would produce an overestimated computation of the cast-in-place pile capacity. In addition, the low specific weight suggests that the effects in the wave equation results of a possible nonlinear stress-strain relationship of the pile material should be investigated. 3. The distinct behavior observed between the two piles in the static load tests was confirmed by the dynamic load tests. Despite their similarity, the pile-soil interaction of the two piles is markedly distinct. The CAPWAP analysis suggest that the skin friction of the CFA-2 pile is higher than the CFA-1 one. On the basis of the current soil investigation performed, the higher capacity of the CFA-2 pile docs not scem to be explained only by a stronger soil around the shaft or at the pile point. The hypothesis of the influence of the grout pressure on the pile-soil interaction cannot be disregarded. Also, since a sand lens intercepts the CFA-2 pilc just above the tip, a distinct response of this soil to the installation procedure might affect the pile end bearing. 4. For the two piles, the maximum mobilized capacities in the dynamic tests ( 1 1 14 and 21 82 kN) are in reasonable agreement with the results of the static load tests (1006 and 1473 kN). The higher CAPWAP capacities are due to the higher displacements achieved in the dynamic tests. A comparison between static and dynamic load tests performed on the same pile should account for the preloading effects. Cast-in-place piles may develop increasing
Figure 5. Static load tests and CAPWAP static simulations for the CFA-1 (a) and the CFA-2 (b) piles.
point resistance over to large displacements. If the soil resistance degradation is reduced, it is expected that the succession of static or dynamic cycles of loading will lead to increased pile capacities. 5 . In order to compare the results of the static and dynamic load tests, the permanent displacement at the end of a cycle of loading should be added to the next one. As shown in Figure 5 , the CAPWAP static simulations shows a remarkable agreement with the trend of the static load tests. 6. The results here presented demonstrate the potential of the high strain dynamic testing as a valuable tool for the assessment of the load-settlement behavior of drilled shafts. ACKNOWLEDGMENTS The authors would like to express their gratitude to Eng. Alexandre Chwist (Estacas Premold), Eng. Janies Barossi (Sondar) and Norberto E. Calliari for their support in performing the tests here reported. REFERENCES Aoki, N. 1989. A new dynamic load test concept. Proc. Discussion Session 14, TC Pile DiYvirig, X I 1 lilt. Coi$ Soil Mech. Fouiidcition Erig., Rio de Janeiro: 1-4.
714
Aoki, N. & V.F.B. de Mello 1992. Dynamic loading test curves. Proc. 4th Itit. Cot$ Applic. of Stress- Wuve Theory to Piles: 525-530. Beim, J. & N.Aoki 1996. Dynamic load test method with variable energy. Proc. 5th lilt. Conf: Applic. of Stress- Wave Tlieoiy to Piles: 274-28 1. Bernardes, G. P. 1989. Dyiiciiiiic mid stcitic testbig of large model piles in sciiid. D.Sc. Thesis, Norwegian Institute of Technology, Trondheim. Chamecki, P.R., A.C.M. Kormann, N.A. Nascimento & A.S. Dyniinski 1998. Sitio experimental de geotecnia da UFPR objetivos e dados preliminares. Proc. of X I COBRAMSEG, ABMS, Bmsilia, Brazil: 8 19-826. de Mello, L.G. & S. Paraiso 1998. Variable energy dynamic load test on 1.0 m diameter CFA pile. BAP 111, Belgiuii: 32 1-334. Dkcourt, L. 1998. Ruptura de fundaG6es e coeficientes de segur a n p a luz do conceit0 de rigidez. Proc. o f X I COBRAMSEC, ABMS, Brcisilia, Brazil: 1599-1606. Hussein, M., G. Likins & F. Rausche 1996. Selection of a hammer for high-strain dynamic testing of cast-in-placeshafts. Proc. 5th 1nt. Coi$ Applic. of Stress- Wave Tlzeory to Piles: 759-772. Kormann, A.C.M. 1999. Comportamento de argilas rijas: aspectos geotkcnicos da Forma@io Guabirotuba. A m i s du M e w Redoiidci Ccimcteristiceis Geote'cniccis du Fosrii~ip7o Gu(ibirotubLi,ABMS/UFPR, Curitiba, Briizil: 1 19- 128. Korniann, A.C.M., P.R. Chamecki, N.A. Nascimento & A.S. Dyminski 1999. Load tests on continuous flight auger piles in the Guabirotuba Formation. Proc. X I Punciiiiericciti Cot$ on Soil Mecli. iirid Geotech. Eiig. : 1537- 1544. Kormann, A.C.M., P.R. Chamecki, L. Russo Neto, L. Antoniutti Neto & G.P. Bernardes 2000. Estacas hklice continua em argila sobreadensada: comportamento em provas de carga est6ticas e dinbmicas. Proc. SEFE l V , S3o Paulo, Brazil. Likins, G.E. & M.H. Hussein 1995. High strain dynamic testing of drilled shafts and cast-in-place piles. Deep Foulid. liist., 20th Aiiriuul Menibera Cot$ Meet., Cliurlestoii. Liu, C., Q. Lin & F. Shi 1996. Determining the bearing capacity of large-diameter bored cast-in-situ piles by high strain dynamic pile-testing. Proc. 5th lrit. Coi$ Applic. of StressWuve Tlieoiy to Piles: 797-804. Maiorano, R.M.S., C. Viggiani & M. Randolph 1996. Residual stress system arising from different methods of pile installation. Proc. 5th lilt. Coilf: Applic, of'Stress-Wuve Theory to Piles: 5 18-528. Massad, F. 1992. Sobre a interpretaq5,o de provas de carga em estacas, considerando as cargas residuais na ponta e a revers3o do atrito lateral. Park I: solos relativamente homogEneos. Solos e Roclzas 15 (2): 103-115. Niyania, S. & N. Aoki 1991. Correla@o entre provas dc carga dinbmica e esthtica no campo experimental da EPUSP/ABEF. 2'. Seni. de Erig. de FuridugGes EspeciLiis, SCO Ptiiilo, B t - ~ i ~ i285-293. l: PDI 1999. Pile Dt-iviiig Aiidyzer, PAL Model, Users Mariual. Cleveland: Pile Dynamics. Rausche, F., B. Richardson & G. Likins 1996. Multiple blow CAPWAP analysis of pile dynamic records. PI-oc. 5th bit. Cor!f:Applic. of Stress-Wcive Theor? to Piles: 435-446. Salamuni, E. 1998. Tectdriicci dci bucia Sediiiientar de Curitibu (PI?). P1i.D. Thesis, Instituto de GeociEncias e CiEncias Exatas - UNESP. Rio Claro. Brazil. Seidel, J. & F. Rausche 1984. Correlation of static and dynamic pile tests on large diameter drilled shafts. Proc. 2nd Iiit. Conf;Applic. of Stress- Wcive Theory to Piles: 3 13-3 18. Seitz, J.M. 1984. Dynamic testing of bored piles i n noncohesive soils. Proc. 211d lrit. Cot$ Applic. of Stress- Wuve Tlieorj to Piles: 201 -209.
Application of Stress-Wave Theory to Piles, Niyama & Beim (eds) 02000 Balkema, Rotterdam, ISBN 90 5809 150 3
Static and dynamic testing of the 'Campile' - A displacement, cast-in-situ pile David J. Klingberg & Phi1 Mackenzie Wagstaff Piling Pty Limited, Brisbane, Qld, Australia
ABSTRACT: Throughout the world, developments are continually being made in the foundation industry to overcome construction and environmental difficulties. The need for a low noise and vibration method for installing piles led to the development of such a system called the 'Campile'. The 'Campile' was developed in Australia by Wagstaff Piling as a means of installing piles on sites that are sensitive to noise and vibration, whilst maintaining the advantages of a displacement pile. This paper outlines the results of a dynamic and static test program conducted on a 'Campile', with the static load test taken to failure. The test results indicate a good correlation between the two test methods. the probe is replaced by concrete or grout as the probe is removed from the ground. The finished pile may be reinforced either partially or full-length, depending on the design requirements. The Campiles installed for this project were at a developmental stage and were constructed as a straight-shafted, nominal 450mm diameter pile. Further advances have since been made with the design and installation of the Campile which allows for other sizes and profiles to be constructed, including a 'screw-shaped' pile which provides increased shaft resistance and greater overall pile capacity.
1 INTRODUCTION
Within the foundation construction industry, environmental considerations with respect to noise and vibration levels have become of more concern in recent years. Modern piling contractors continually look for new ways to provide cost-effective and technically sound solutions to these environmentally challenging projects. The Campile was developed by Wagstaff Piling, Australia as the result of such a need. The Campile is a displacement, cast-in-situ pile that is installed under low noise and vibration conditions. As a displacement pile, it combines the advantages of non-removal of soil, but maintains the low noise and vibration of traditional non-displacement, drilled piling methods. This paper presents the outline of the test pile programme completed on a Campile, summarises the results of the 'Class A' dynamic test data and provides a comparison of these results with the static test data. Some recommendations are also included in this paper with respect to methods to be adopted for wave equation analysis for displacement, cast-in-situ piles.
3 SITE GEOLOGY
The project for the installation of the Campiles was located at an industrial site near Sydney, Australia. The site was underlain by loose to medium dense sand and silty sand layers to depths of approximately 4 to 5 metres over a dense sand layer to 11 metres. The target penetration for the Campiles ranged from 5.5 to 7.5 metres where the base resistance was expected to be high in the dense sand layer.
2 DESCRIPTION OF THE CAMPILE
4 TEST PILE PROGRAMME
The Campile is a proprietary pile of Wagstaff Piling and is constructed by screwing a specially designed probe into the ground, which displaces the soil as the probe penetrates the ground. The void created by
4. I Dynamic test results
Test pile TP was installed as a sacrificial pile for the project. The pile was a 450mm diameter (nominal)
715
Campile with a total length of 7.0 metres and a drilled depth of 5.8 metres. The Campile was constructed with grout using a 40MPa design strength. Test cubes were crushed prior to the dynamic test being performed and indicated a strength of approximately 32MPa at 7 days. The dynamic test was conducted when the pile was 8 days old. To enable the test piles to be subjected to dynamic and static testing, a special 40 MPa concrete head was cast onto the pile to contain the impact and bursting stresses from the impact of the test hammer. The test hammer was a five (5) tonne drop hammer, with drop heights up to 13OOmm used during the testing. The dynamic testing and CAPWAP@analysis were completed as ‘Class A’ predictions (ie. without knowledge of the static test results). The typical dynamic test data is shown in Figure 1, where the difference in impedance between the pile head and pile shaft (as expected) is clearly seen. The results of the dynamic testing are summarised in Table 1, with the results of the CAPWAP@ analysis included in Table 2. PILETEST
Table 2. ‘Class A’ CAPWAP@analysis summary. Pile ShaR End Total No. Resistance Resistance Resistance
TP
Lrz
50 m s
...... .................. ....... ..... .........
0 ..................
F
(W 1,797
Table 3. ISL compression test results summary. Load Pile head Creep Creep (kN) Deflection Correction Corrected Deflection (mm) (mm) (mm) 225 0.52 0.00 0.52 450 1.63 0.00 1.63 675 3.06 0.03 3.03 900 4.75 0.20 4.55 1,125 8.15 0.27 7.88 1,350 17.70 5.89 11.81 1,575 25.60 6.41 19.19 1,800 50.00 10.27 39.73
51 DROP W E R
3.2mS
(W
1,427
The test pile was subjected to an incremental sustained load (ISL) compression test in accordance with the requirements of the Australian Standard@ Piling Code - AS2159 (1995). The load test was conducted to a load level where plunging failure of the pile occurred. The results of the ISL load test have been summarised in Table 3 and shown graphically in Figure 2. As dynamic load tests cannot predict any time dependent deflections and to enable a realistic comparison to be made, all deflections from the ISL test have also been corrected for creep. The creep corrected values are given in Table 3.
PD(PLEDRMV,UULIZER
RMX 1890 kN EMX 43.30 kKm DMX 22.8mn WU2 100kN LE 6.10111
370
4.2 Static test results
TP-RST
.....................
(W
1‘
4.3 Comparison of static and dynamic test results
kN
~
l ..,. p..
.. ... .
.
.x . l I
The results of the static and dynamic tests have been plotted in Figure 3. It can be seen that the CAPWN@ analysis provided a conservative load-deflection behaviour compared to the static test data. Afker the results of the static test were known and compared to the ‘Class A’ predictions, the CAPWAP@analyses were re-analysed to provide a predicted deflection behaviour that more closely modelled the static behaviour. The results of the re-analysis are given in Table 4. It is noted that the revised CAPWAP@analysis did not alter the ultimate capacity of the pile, but adopted a different distribution of capacity to achieve the ‘better match’ to the ISL test results. The comparison between the revised CAPWAP@analysis and the static test data is given in Figure 4 and Table 5 . The revised CAPWAP@ analysis showed excellent agreement with the ISL test results.
............................................................ & ; s
.. ,. I
I
I
Figure 1. Dynamic test data for test pile TP Table 1. Pile No. TP
Dynamic testing results summary. EMX RMX TC Set Stroke (mrn) (mm) (mm) (kNm) (kN) 6.8 43.3 1,890 1,300 16.0
Set = pile set in mm/blow TC = temporary compression in mm EMX = max. energy transferred to pile head in kNm RMX = PDA capacity estimate in kN 716
Figure 4. ‘Revised’ CAPWAP@ and compression test results.
ISL
4.4 Discussion on results
Figure 3 Comparison of ISL compression test and ‘Class A’ dynamic test results. Table 4. ‘Revised’ CAPWAP@analysis summary. Pile Shaft End Total No. Resistance Resistance Resistance TP
(W
(lm
776
1,021
Table 5. Revised CAPWAP@and ISL test results summary. Load ‘Class A’ Revised (kN) CAPWAP@ CAPWAP@ Deflection Deflection (mm) (mm) 225 1.78 0.48 450 3.55 0.96 675 6.34 1.50 900 10.59 4.39 1,125 14.89 10.67 16.95 19.20 1,350 1,575 26.23 23.22 43.85 43.79 1,800
(W 1,797
compression ISL Creep Corrected Deflection (mm) 0.52 1.63 3.03 4.55 7.88 11.81 19.19 39.73
The re-analysis of the dynamic test data and the corresponding better match with the ISL test results indicated that the shaft resistance on the pile was significantly higher than initially obtained from the ‘Class A’ CAPWAP@ analysis. The revised CAPWAP@showed approximately twice the shaR resistance of the ‘Class A’ analysis. The re-analysis was achieved with significantly lower skin damping and quake parameters, with only minor changes to the toe damping and quake parameters. It should be noted that shaft radiation damping parameters were used in both the ‘Class A’ and ‘revised’ CAPWAP@analyses (CAPWAP@Manual, 1996). Without the use of this modelling tool, the predicted ultimate capacity from the CAPWAF@ analyses may have been significantly lower than the static test value (approx. 15% to 20% lower). A comparison of the shaft resistance values from the ‘Class A’ and ‘revised’ CAPWAP@analyses was also made with standard pile design calculations using the borehole information for the site (Poulos and Davis, 1980). Three separate design calculations were performed on the borehole data: (a) bored pile design; (b) grout-injected pile design; and (c) driveddisplacement pile design. A summary of these design calculations is given in Table 6. It can be seen from Table 6 that the ‘Class A’ values from CAPWAP@correlated with bored pile design parameters, whereas the revised CAPWAP@ values were slightly higher than those from a driveddisplacement pile design. This confirmed that the Campile was performing in the manner in which it had been designed - that is, as a displacement pile and not as a non-displacement pile.
717
Table 6. Design calculations summary. Depth Stratum Bored Pile (m) Design (kN) 2.5 L/MD Sand 107 3.2 Loose Sand 29 4.5 MD Sand 73 5.8 Dense Sand 199 Total 408
GIP Design (kN) 193 52 132 22 1 598
Driven Design (kN) 214 58 147 23 8 657
CAPWM@
(W
‘Class A’ 126 45 61 138 370
‘Revised’ 253 84 185 254 776
5 CONCLUSION
REFERENCES
The Campile is a proprietary pile type developed by Wagstaff Piling, Australia as a low noise, low vibration, cast-in-situ, displacement pile that produces limited spoil. The Campile was developed as part of an ongoing research programme to provide piling solutions for environmentally sensitive sites, particularly inner city developments where noise and vibration considerations are paramount. The test pile programme for the Campile installed for this case study showed that dynamic testing methods could be used to predict the behaviour of the Campile under static loading conditions. The ‘Class A’ CAPWAP@analysis provided a conservative prediction of the pile behaviour. A ‘revised’ CAPWAP@ completed after the static results were known showed excellent agreement between the dynamic and static methods. It was noted that the revised analysis provided a significantly larger shaft resistance than the initial ‘Class A‘ analysis and adopted significantly lower shaft damping and quake parameters. Comparison of the shaft parameters obtained from the CAPWAP@analyses with calculations from the borehole information for the site indicated that the Campile was performing as a displacement pile. That is, the shaft parameters obtained from the test programme on the Campile were similar to those expected for a driven, displacement pile. It was also noted that the Campile installed for this case study was a development stage of the pile type and was installed as a straight-shafted pile. Further advances in the pile installation process have since been made with the Campile which allows for various sizes and profiles to be constructed, including a ‘screw-shaped’ pile which provides increased shaft resistance and greater overall pile capacity. During the installation of the Campile, torque measurements were also recorded but have not been included in this paper. However, it is noted that the torque readings show good correlation with the CPT data collected for the site.
Australian Standard Piling Code. AS 159 (1995). Standards Australia. CAPWAP@ Manual (1996). Pile Dynamics Inc. Cleveland, Ohio. 1980. Pile Poulos, H.G. and Davis, E.H. foundation analysis and design. John Wiley and Sons.
71a
Application of Stress-Wave Theory to Piles, Niyama & Beim (eds) 02000 Balkema, Rotterdam, ISBN 90 5809 1503
LT the final word? Correlation between DLT and SLT H.Goldemberg & J. J.Goldemberg Geotecnica Cientec, Buenos A i m , Argentina
ABSTRACT: Due to the increase in service loads of foundations, a world wide trend is being followed by geotechnical engineers and designers to measure the real behaviour of the pile-soil interaction. Historically, this was defined by the Load-Displacement Curve, obtained from a Static Load Test, and the concept of bearing capacity.. .if those involved in the project arrived to an agreement on its definition. Technology developed and the concept of Dynamic Load Test (DLT) arrived, bringing with it a much economic and faster way of measuring the response of piles to imposed loads. But is it this what the geotechnical engineer requires? A comparison between DLT and SLT is analysed throughout the eyes of a geotechnical engineer and not from the perspective of a testing house. Different types of piles were tested with both methods, seeking for correlation, side effects and installation influence in results, aiming to know the real performance vs. predictions. 1 INTRODUCTION In the first half of the ‘SOS, it was introduced in Argentina the pile testing speciality based on the Stress Wave Theory; at the beginning throughout the Sonic Integrity Test (SIT), as a parameter for control quality of foundations and, afterwards, with the Dynamic Load Test (DLT) in order to measure the pile-soil behaviour. But up to then, how was determined the bearing capacity of a piled foundation? The answer is simple, in the same way than in the rest of the world, that is, with a Static Load Test (SLT). As we all know, in spite the familiarity SLT has within the engineering community.. .that, if we first agree on what type of SLT, due that the ASTM D 1143-81 Standard has five different methodologies.. .it is time and money consuming. During those years, where Quality Assurance (QA) started to play an important role in the Argentinean construction market, as well as the increment in service loads, leaded to an increase in the demand for verifying piling works. As well, a reactivation of the local economy, the availability of new technology in construction methods and tighter schedules for finishing site works, allowed the DLT to be introduced as an alternative to the cumbersome static test.
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It is the objective of this Paper to present correlations between Dynamic and Static Load Tests performed on the same pile, including all the information that the Geotechnical Consultant had at the moment of analysing the results and highlighting how it was arrived to those conclusions. It is the moment to mention that in all cases the Dynamic Load Tests were performed using the FPDS-3 equipment, developed by TNO Building & Construction Research (the Netherlands), while the SLTs were conducted under ASTM Standard. When the Static test was carried out, the transfer load structure assembled consisted in a beam and reaction piles anchored well below the influence area of the test pile. 2 CASE HISTORIES. 2.1 Buildings, The site was a group of buildings constructed in the City of Buenos Aires, Argentina. The ground conditions are described in the following geotechnical profile (Figure l), and the pile type was of substitution, reinforced concrete and casted in-situ under bentonite mud (pile characteristics are presented on Table 1).
Table 1. Pile characteristics. Pile Bentonitemud Type I J J I1 111
J
Diameter (m) 0.35 0.40 0.50
‘slow method’ and then 7% of the foundation population (29 piles) were going to be tested with the Dynamic Load Test (the tests information is presented on Table 2) Conclusions: The results of the homologated tests are presented on Figure 2, where it is clearly seen the good correlation between DLT and SLT. The quality of this match depends dramatically on the quality of the geothecnical investigation, the knowledge of the mechanical properties of the concrete and the concrete consumption, in order to check variations in the cross section of the piles.
Length (m) 14.0 15.0 15.0
2.2 Industrial.facility. The site consisted in the expansion of an industrial facility in the South Area of Gran Buenos Aires, where 250 precast reinforced concrete piles were going to be driven with a diesel hammer. Pile data is available on Table 3 . In order to determine the pile length in different areas of the building site, Pile Driving Analysis (PDA) was performed, monitoring Driving Resistances, Enthru Energy, Stresses, etc. The 100% of the pile population was tested with SIT searching for cracked piles, 10% of the foundation (20 piles) was tested dynamically and only 1 ‘slow’ SLT was performed to have a correlation between both methods. The summary of these investigations is presented on Table 4.
Figure 1. Geotechnical profile.
Table 2. Test programme* Test Pile Quantity Service Max. Test Maximum Type Type Load Load Settlement (MN) (MN) (mm) DLT I 10 0.6 1.4 2.6 I1 10 0.7 1.9 3 .O I11 9 0.9 2.4 3.9 SLT I 1 0.6 1.2 2.2 I1 1 0.7 1.4 2.1 Ill 1 0.9 1.8 2.6 Note *: information is presented only for those piles where DLT and SLT were performed.
Figure 2. DLT and SLT load-settlement curves. Table 3. Pile characteristics. Pile Cross section Type (mxm) Concrete precast 0.4x0.4
The foundation project consisted in 408 piles placed under caps in groups of one, two or three units each. All the installed piles had to be controlled through the Sonic Integrity Test (SIT) in order to verify that they were fiee of damages and to select which ones were to be subject for DLT and SLT. It was decided, by the Engineer, that the first three piles were going to be tested statically with the
Table 4. Test programme. Test Quantity Service Type Load (MN) PDA 40 DLT 20 1.2 SI.T 1 1.2
720
Length ( 4 19.0
Max. Test Load (MN)
Maximum Settlement (mm)
2.9 2.4
4.8 3.8
It is interesting to discuss what happened with this correlation. Due to the unreal short period of time allowed to finish all the driving works, the Engineer decided to perform the Dynamic Load Test immediately after driving- although GEOTECNICA CIENTEC intended to persuade him not to do so because a less total resistance was going to be obtained due to the influence of pore pressure. The geotechnical information, described on Figure 3, shows the stratigraphic profile throughout SPT and CPT tests. Because discrepancies were detected during the original borings in different areas of the site, Dutch Cone was used. It was also very useful the Pile Driving Prediction (PDP), performed before the first pile was driven in order to select the most suitable diesel hammer. On Figure 4, the results of the load tests performed on pile No 135 are plotted, where the first DLT carried out did not fidfilled the expectationsdue to what was mentioned above. After the results were available, the Engineer decided to perform the SLT two weeks later, having in mind that the test could not be taken to failure because it was a service pile. Five days after finishing the Static test, the piling rig was used to redrive pile No 135 to monitor its Dynamic behaviour. As it can be clearly observed, the differences between the first DLT and the SLT as well as among the Dynamic ones was caused because it was not taken into account that in cohesive saturated soils, the driving effects on remoulding soil layers and the slow dissipating pore pressure cause a reduction of the bearing capacity.
Conclusions: Dynamic Load Tests are affected by local geotechnical characteristics as well as by soil conditions at the moment of carrying out the measurements. This is the reason why the frst DLT did not provide reliable information, because the behaviour of the pile-soil interaction during testing was different than that at the design stage. It is the authors’ opinion that in order to perform a better geotechnical design of piles, more efficient and foundation works less expensive, it must be to carry out CPTs tests during the soil investigation programme. In this way better Signal Matches will be obtained between calculated and measured Upper Travelling Waves, because more exact soil models can be evaluated.
Figure 4. DLT and SLT load-settlement curves.
2.3 Bridge.
In the Province of Santa Fe, Argentina, due to yearly over floods of the Parana River- caused mainly by heavy rains during the rainy season, many bridges in the area of the lowlands had to be rebuilt and some others expanded. In this case, the project consisted in raising the level of the road as well as increasing the length of the bridge fi-om one span 10m long to a deck of two spans 15m each. The geotechnical design for the foundation was based on the soil strata illustrated on Figure 5 , adopting bored piles casted under bentonite mud. The pile details are included on Table 5. The North, South and Intermediate piers had three piles each, which were tested with SIT, aiming to detect defects such us cracks, soil inclusions, contaminated concrete and their real lengths. The testing programme carried out is shown on Table 6. Table 5. Pile characteristics. Pile Bentonite mud Type Bored J Figure 3. Geotechnical profile.
721
Diameter (m) 1.20
Length (m> 21.5
Figure 6. DLT and SLT load-settlement curves.
Figure 5. Geotechnical profile. Table 6. Test programme. Test Quantity Service Max. Test Maximum Type Load Load Settlement (MN) (MN) (mm> DLT 1* 4.0 6.5 7.3 SLT 1 4.0 7.2 8.6 Note *: Originally 1 SLT and 1 DLT were projected, but as one pile was damaged it was agreed to perform a second DLT.
From the interpretation of the Sonic Integrity Tests signals, it was concluded that pile NO4 had an impedance reduction. Afier a detailed analysis using the Signal Matching technique with TNOWAVE, it was calculated that the abnormality consisted in a reduction of cross section from 1.20m of nominal diameter to 1.12m extended over a length of 0.5m and at a depth of 14.6m. With this result, it was decided to go further on in the investigations, in order to determine the real behaviour of the pile-soil interaction and measure if the foundation element was suitable of transferring loads to the surrounding stratigraphy. A Dynamic and a Static Load Tests were going to be conducted on pile N04.
Due to the local geological conditions, mainly sandy soil, a ‘quick’ Static Load Test was going to be performed because the creep effect was not critical for the bridge behaviour. The results are plotted on Figure 6 jointly with its dynamic counterpart. Conclusions: As it is clearly presented, the Dynamic Load Test can have an accurate static correlation, inclusive for high loads- 7MN as in this case, As in the previous cases, of the building and the industrial facility projects, it is an irrevocable condition - in order to obtain a good match or correlation between dynamic and static measurements of a same phenomenon pile-soil interaction - to have an exhaustive geotechnical investigation, knowledge of the parameters of the materials used for the construction on the foundation and the pile shape. Although some skepticism, that still over flies on some Engineer’s minds, Dynamic Load Testing is reliable if the above conditions are fulfilled, inclusive for high loads in cast in-situ piles. 3 GENERAL CONCLUSIONS. In the above case histories, where piles varied in dimensions and construction methods- fi-om driven precast concrete to bored piles under slurry, throughout Franki type; installed in clay, lime and sand affecting the pile-soil interaction in different ways under small to big loads, they all have a commo n den0 minator. That common denominator is, leaving aside the fact that all tests were properly performed, recorded and post-processed, the quality of the basic inforrnation. In other words, the geotechnical information, dimensions and pile shape, mechanical properties of materials used and the installation procedure. For both types of test, Dynamic (DLT) and Static (SLT). that information is vital. Notable contradiction, because it is required for both designing the test and interpretation of results. It is useless to obtain
722
huge amount of data, in white paper in case of a SLT or a computer file for a DLT, in order to be plotted as the ‘Load-Settlement Curve’ and handed in to the Client as a great achievement certifying that ‘the pile will stand the load’ without the geotechnicalstructural diagnosis. In other words, the interpretation! If not, effort, time and money would have been wrongly spent.. .three not rehndable goods in any aspect of life. But making an abstraction and travelling to the Contractors’ Paradise, where no testing is required and QA has not been invented- not even as a concept. Nevertheless, would not be necessary to know the geotechnical investigation, material properties and selected pile in order to construct the project? In almost all the above case histories, a good correlation between Dynamic and Static Load Tests was obtained, satisfying the expectations for homology of both methods. It is a fact the acceptance for DLT fiom the local market and its growing reliability, due that it provides similar results than the Static Test with almost no delays on site and reduced budget. 4 REFERENCES. ASTM, D 1143-81 (Reapproved 1994). “Standard Test Method for Piles Under Static Axial Compressive Load”. ASTM ASTM, D 4945-89. “Standard Test Method for High-Strain Dynamic Testing of Piles”. ASTM C. R. Mullins (1 992). ‘‘Large diameter pile test project”. Proceedings of the Conference on Recent Larfe Scale Fully Instrumented Piles in Clay, London, June 1992 Carol1 L. Crowther (1988). “Load Testing of Deep Foundations”. German Society for Static and Dynamic Pile Tests (1997). “Recommendations for Static and Dynamic Pile Tests”
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Application of Stress-Wave Theory to Piles, Niyama & Beim (eds) 02000 Balkema, Rotterdam, ISBN 90 5809 150 3
Results of an international pile dynamic testing prediction event A. Holeyman Universitk Catolique de Louvain, Louvain-la-Neuve, Belgium
J. Maertens Katholieke UniversiteitLeuven, Leuven, Belgium
N.Huybrechts & C. Legrand Belgian Building Research Institute, Brussels, Belgium
ABSTRACT: An international prediction event was carried out within the framework of a 30-pile testing program organized in Belgium. That program called upon several testing methods: static load tests, Statnamic testing, and dynamic testing. This paper provides a summary of received predictions and results obtained from the static pile load tests, which were carried up to failure of the instrumented piles. The comparison between predictions is made using load-settlement curves with reference to the results of the static load tests. A companion paper reports on the project background information that was required to prepare the prediction event, including a description of the pile types and results of an extensive soil investigation program. al, 1999b). Predictions could be established based on the dynamic load tests, the geotechnical investigation, experience, or a combination of the above. The piles types, layout, and the site investigation are described in a companion paper (Holeyman et al, 2000). The present paper focuses on load tests results and their comparison with results of the static load tests.
1 INTRODUCTION 1.1 Program Background
A national research project has been organized by the Belgian Building Research Institute in order to establish the performance of different types of castin-place ground displacement screwed piles. The program included the installation and testing of 30 test piles that allowed the organization of a prediction event. That prediction effort was undertaken with the hope to document the profession's ability to estimate these new piles behavior based on standard investigation means as well on dynamic testing. Six different types of ground displacement piles were installed (five of each) and tested: one prefab and five cast-in-place screwed types: Atlas, De Waal, Fundex, Olivier, and Omega. An extensive soil investigation was performed as part of the research project, including in situ tests (CPT, PMT, SPT, DMT, SASW) and laboratory tests on undisturbed samples. 1.2 Prediction preparation A project synopsis had been prepared to invite interested parties to make those predictions (Holeyman et al, 1999a). It included a description of the pile types, site characterization, the static load test procedure, the dynamic load test procedure and the format of the prediction. Interested predictors were sent the complete information, available as laboratory and in situ investigation and dynamic load test results (Holeyman et
2 DYNAMIC LOAD TESTS 2.1 Procedure for dynamic load tests The loading device used to impact the 12 piles installed for that purpose was a 4 tons drop hammer operated by a crane. A sequence of several blows was applied to each pile. The drop height sequence most often applied was as follows: 0.40m, 0.80ni, 1.2m, 0.8m, and 1.2 m. Dynamic measurements of strain and acceleration were acquired for all 12 piles using a TNO FPDSS system. In addition 6 piles were also monitored using a PDI PDA-PAK system (Prefab, Fundex, and De Waal piles) and one Omega pile was monitored using a PDI PAL system. A 0.4m-diameter head was cast on July 6'' on top of the 10 cast-in-place piles. The transducers were attached generally 0.8 m from the top of the approximately 1.5 m high head. Displacements were acquired using a laser system. 2.2 Distribution of dynamic load test measurements The results of dynamic load test measurements (pile head force, velocity and displacement) were made 725
Figure l a : Long Piles Load Settlement Curves - Prefab, Fundex and De Waal
726
Figure 1b : Long Piles Load Settlement Curves - Olivier, Omega and Atlas
727
available under a digital format to parties that had expressed an interest to make a prediction on that basis. Also characteristics needed to interpret the measurements were provided in part in Table 1 of the companion paper (in particular dimensions and properties of the pile heads extension) and in part in the files containing the measurements. Additional characteristics (wave propagation speed, density, pile impedance, etc.) needed to further analyze the measurements were distributed together with the measurements. Interested parties obtained the digital files of the events by e-mail, which required the structuring of a vast amount of information, totaling more than 8 Mbytes of digital records. For each pile type a directory was established (for example the Prefab directory), containing subdirectories according for each pile of that specific pile type. The ‘pile number’ subdirectory (for example subdirectory pile A7 in the directory Prefab) contained a word file: (e.g. ‘A7info.doc’) and was further subdivided into the following subdirectories: TNO files, ASCII files, PDA files, and Displacement files. The word file ‘A7info.doc’ gave supplementary information about the dynamic load test on the pile A7 (the blow numbers, the drop height, and field notes). It was thus possible for the predictors to reprocess the raw signals using adjusted pile parameters, number of samples, etc. 2.3 Meas urements nomin a I interpretat ion
The choice of the relevant moduli and sections is often considered as part of the predictors’ art and was purposely left open to some degree, as is usually the case for cast-in place piles. It was emphasized that all files had been uniformly acquired using a nominal modulus of approximately 40,000 MPa and a nominal wave speed of 4,000 m/s at the measurement section (i.e. in the concrete of the cast head). The pile heads were cylinders with a diameter of 0.4 m, except for the Prefab pile where the current 0.35x0.35m section prevailed. It was the predictor’s responsibility to assess the measurement section modulus adequate for his prediction. It was also emphasized that the concrete of the tested pile below the added head had a different modulus and a different section. It was also the predictor’s responsibility to assess the appropriate section and modulus for the shaft. Peripheral information allowing the predictor to perform that important assessment included Table 1 of the companion paper, the digital signal themselves (e.g. impedance match or 2L/c check), integrity tests of the piles. Strength and ultrasonic wave speed measurements on concrete samples cast at the time of installation of the piles and concreting of the pile heads were made available finally. Low strain testing had been performed on all piles both the
BBRI and CEBTP and results provided to predictors. 3 RECIEVED PREDICTIONS 3.1 Reporting format The predictions are reported herein under an anonymous format in order not to stigmatize those with less accurate predictions. Each prediction is however labeled with a code corresponding to the prediction type. Each predictor is thus enabled to position his own prediction within the cluster of results and encouraged publishing his prediction procedure, using the present paper as a reference. 3.2 Predictions types Results from 10 predictors had been received on November S’’ 1999, the ultimate submittal date. Contractors had also predicted the ultimate bearing capacity of their own piles (they were not asked to supply the load-settlement curves.) According to the reference data used to cast those predictions, the following labels have been used: - “CPT” for predictors using the CPT results. - “PMT” for predictors using the PMT results. - “LAB” for predictors using the laboratory results. - “DLT” for predictors using the Dynamic Load Test results. - “STN” for predictors using the Statnamic Test results. The CPT predictors used different methods, including ultimate state design as well as load transfer curves. All the contractors’ predictions were made using CPT results and De Beer’s 1974 method. The PMT predictors used the pressiometric approach that provides stress-displacement relationships for the shaft and the base. The LAB predictor used a loadtransfer functions method based on plasticity indices. The DLT predictors’ methods included either CAPWAP or SIMBAT: the soil parameters in a model are adjusted to get the best match between the measured and the predicted signals of a Dynamic Load Test. SIMBAT is an empirical method converting the dynamic reaction to a static reaction. The STN predictor used the Unloading Point Method (UPM) to predict the static load test. It was mentioned by the predictor that, due to strain rate sensitivity of clayey soils, a 30% reduction coeffcient had to be applied on the usual UPM method. A hyperbolic approximation of that reduced function was then calculated. This is the reason why those predictions are labeled as “0.7 STN”. It should be noted that the STN predictor was not provided with the results of the dynamic load tests, and that no TNO-WAVE prediction was submitted.
728
3.3 Predictions classes Predictions had to be made before static pile load tests were performed in order to qualify as Class A type predictions. If predictions were made after the static pile load tests, they qualified as C type predictions, according to accepted definitions of predictions classes (Lambe, 1973). Each prediction for each pile can be classified according to its submittal date relative to the date of static loading. Table 1 shows that most of the predictions are Class C. The only Class A predictors were CPT 1 (except for piles A1 & A4) and Contractors Atlas and Fundex. Other Class A predictions were those of predictor DLT2 for A3 pile and of predictor PMTl for C2 to C4 piles. 4 STATIC LOAD TESTS 4.1 Procedurefor static load tests The static pile load tests were to comply with the following loading guidelines, referring to Q,,,, the maximum anticipated test load, chosen with the hope to cause bearing failure: - A pre-load stage of maximum 5% of Q,,, was applied in order to check the measurement equipment and the centricity of the applied force, - 10 maintained load steps with equal AQ until Q reaches Q,,, - No intermediate unloading cycles Duration of maintained load step of 60 minutes - Load test was performed until a pile head settlement 2 15% Obase was reached - When the pile head settlement has reached a value of 25 mm, subsequent load steps can be applied using a smaller increment (AQ/2), in order to refine the pile load-settlement curve as it approaches failure, - Unloading in 5 steps of 10 min. each, except for final unloading (30 min at least of monitoring). The system used to apply the maintained loads on the piles called upon a sophisticated hydraulic regulation that guaranteed a tolerance of 5 kN. That system had just been developed by the BBRI. The 3 MN reaction was provided by a kentledge consisting of concrete blocks. Besides load and settlement monitoring, extensometers provided longitudinal strains along 5 to 7 shaft segments along the pile length. The results provided by those more detailed measurements are to be reported elsewhere. Such a procedure requires a value for the ultimate capacity R, of each pile. Those capacities were estimated by the BBRI and the national experts using De Beer’s method based on the CPT tests results (De Beer, 1974). The load increments A Q were actually: - R,,/8 for Atlas, Fundex, Prefab and Olivier piles. - RJ0 for De Waal and Omega piles.
The ultimate capacity was considered reached when the pile head settlement was equal to 10% Obase.It should be noted that a maximum Constant Rate of Penetration ( C W ) of 0.6 m d m i n was enforced towards the end of the loading procedure for all piles (except for piles A1 and A4). 4.2 Results Figures 1 and 2 show the various load (Q) - settlement (s) curves for the long and the short piles, respectively. “SLT” refers to the Static Loading Test. The predictors’ curves are also identified using the labels discussed in Section 2.2. “Contractor” refers to the ultimate capacity predicted by the Contractor. This value is drawn for 30mm < s < 50mm with a bold line. “Target SLT” refers to the ultimate capacity estimated by the BBRI. It is a “box” corresponding - 8AQ (mm) (In)
Shaft Toe 1 STM500 2 DSLSO 3 DSL15
2000 62 0 13651 5468 800 7 3 0 800 1 1 5
4 DSL4S 5 DSLIS
609 37.5 609 3 9 0
6 DSL45
609 35.0
2724
411
20 12 1 465 186 1067 206
Total
Shaft Toe Total
19120 19424 5978 25402 3165 7213 141
960
196 7409 19 1009
Shaft 321 0 02
250
02
220
683
651 3679 1214 1923 1273 1861 570 2131
1320 3850 754 4604 11071 10672 2274 12916
28.2 121 2
9 HYD100
1600 52 0 11662 1312
15971 14501 3822 18326
295 3
17
92.8 90.6 86.3 60.6
1.2 1.5 1.6 12.0
51.0 32.0 82.0 18.0
STM Steam hammer
1735 2862 1480 323 1823 5958 1098 892
4711
8722 1998 13720 >14d
Ilh 100 32.5 1.6 10.0 115h 15 3 1111 250 120 64.6 2.0 16.0 105h
800 11.0 1000 7 6 0
700 508 800 400
29067 6213 35280 >14d
4292
118 4110 >11d
4059 6217
596 4655 31tl 515 6762 >14d
17
7 DSL10 8 DSL72
NYDl00 DSL45 HYD100 HYD65
43h
426
284 735 1259 61 8118 2656
10 11 12 13
Toe Total
1019 3925
4596 1803 7781 1989
918 1843
7860 960 8820 3410 823 4234 11731 4067 15798 1401 774 2176
23.0 5.0 0 1 220
66h 66h
4240 1245 5485
27tl
4200 158 4658 12279 1215 over 13524 22510 3920 26160
29d 90d
21 0 15mi n 27.0 67211 6105 2127 8232 24.0 138h L] - 3822 15.0 15h 10780 3920 14700 13.0 720h 1421 735 2156
70d l4cl l1d lltl 90d
DSL Diesel hammer HYD Hydraulic hammer
were estimated through the wave matching analysis of the dynamic load test signals In the cases of the static load tests, Q, and Cl, were separately estimated when axial forces of the pile were measured, otherwise 0 alone was measured Note that Q is defined as the load corresponding to a pile toe displacement of 10% of the pile diameter The load-displacement curve estimated from the wave matching analysis of the re-driving test signals is compared with the load-displacement curve obtained from the static load test in Figures 13 to 19 for 7 tests out of the 13 comparative tests For the 7 selected cases, set per blow, S, was greater than 1 2mm In cases 1 and 2, S was almost equal to 0 indicating insufficient driving energy to estimate the ultimate pile capacity In fact, the ultimate capacity derived from the re-driving tests for cases 1 and 2 are notably smaller than the ultimate capacity obtained from the static load tests In cases 3, 5 and 9, the time intervals between the end of initial pile driving and the re-driving tests were short (14 hours in cases 3 and 5 , 15 min in case 9), not allowing the completion of the set-up phenomena of the piles In these cases, the ultimate capacity derived from the re-driving tests is clearly smaller than those obtained from the static load tests It is interesting to note that the same hammer (DSL45) was used in cases 3, 4 and 5 and an ultimate capacity of 4 9MN was derived in case 4 745
Therefore, the small values for the ultimate capacity derived from the re-driving tests for cases 3 and 5 can be attributed to the fact that the set-up phenomena have not been completed at the time instants of the re-driving tests In case 8, the ultimate capacity was not obtained in the static load test due to insufficient load capacity of the loading device Therefore. case 8 was excluded from the comparison of the load-displacement curves obtained from the dynamic and static load tests Figure 12 shows the change with time of the setup ratio of each pile after the end of initial pile driving The values of set-up ratio are in a wide range from 1 1 to 10, although the set-up ratio is lower than 4 if the re-driving test is conducted within 3 or 4 days after the end of initial pile driving It may be very difficult to predict the set-up ratio accurately at the current stage of the pile technology Figures 13 to 19 show the load-displacement curves obtained from the static load test and derived from the re-driving test, together with the profiles of the soil layers and the SPT N-values, for piles 4, 6, 7 and 10 to 13 where re-driving tests were conducted at time instants sufficiently after the end of initial driving to permit a full recovery of the pile capacity Various commercial computer programs such as CAPWAPC and TNOWAVE, and KWAVE (developed by Matsumoto & Takei (1991)) were
Figure 14. Load- displacement cunes of case No.6
used for the wave matching analyses to derive the static load-displacement curves. However, the empirical soil model developed by Smith (1 960) was used in all of the computer programs. Even with such conditions, the derived load-displacement curve is in good agreement with the static load test results for each case. While some discrepancy between the derived ultimate bearing
capacity and the measured bearing capacity is seen in some cases, the initial portion of the derived loaddisplacement curve is fairly coincident with the measured curve, suggesting negligible influence of the computer program and the operator Table 3 shows a comparison of the calculated yield load and the yield load obtained in the static load test of each pile. The yield load is defined as the load corresponding to the first rapid increase in the pile displacement on the load-displacement curve. The 746
Table 3. Yield pile capacity derived from the dynamic and static load tests. Static yield capacit! (kN) Static load Matching Case No. error Test Analwis
4
4635 4998 429 1 5880 3822 1 1760 2940
6 7
10 11 12 13
4924 4843 3150 8397 4367 12576 2653
5.9% -3 2 % -3 4% 30.8Y" 12.5% -10 8% 6.5%
Table 4 Equations to estimate soil paraiiieters for the rational soil models for shaft iiiodel k , = 2 7% l ( d ). c, = G/175 8(; 3 2G for base model kl, = ____ . Ch =?r(l- c,)I -> 7T( 1 - v)d '
Poisson's ratio of soil. p, soil dciisitj, cJ pile diameter Shear modulus of soil. T ', shear u a\ e 1clocit> of soil
\c
(;
=pl :2
Figure 20 Ratioiial soil models (after Raiidolph & Siiiions 1986. Raiidolph & Decks 19'92)
difference between the derived yield load and the measured yield is within t10%, except for pile 10 where KWAVE program was employed for the wave matching analysis An error of I-tlOO/b may be acceptable, if the safety factor of 3 is used as is in Japan Even if the accuracy of the dynamic load test is less than the accuracy of the static load test, the increase in the number of tests performed at a site allows an increase of the design value for the bearing capacity (Hayashi et a1 2000)
Figure 21 W a ~ cmatching results n i t h the use or Smith iiiodel
Pile No 10 was re-analyzed by means of KWAVE with the use of the rational soil models developed by Randolph 8L Simon (1986) and Randolph & Deeks (1992) The rational soil models have been incorporated in the KWAVE program In the KWAVE program, the wave propagation in the soil 747
The soil density, p, was assumed to be 1800 kg/m3 for all the soils. In the wave matching analyses, the distribution of zm, and q b alone were assumed. The final matching results using the Smith model and the rational soil models are shown in Figure 21 and Figure 22, respectively. A better matching was obtained in the wave matching analysis using the rational soil models. The distributions of rmx derived in both the wave matching analyses are shown and compared with the static load test results in Figure 23. The corresponding static load-displacement curves are shown in Figure 24, together with the static load test results. It can been seen fiom these figures that the wave matching analysis using the rational soil models results in better predictions of the static load test results. Figure 22. Wave matching results with the use of the rational soil models.
4 CONCLUSIONS This paper reviewed dynamic load tests performed in Japan, showing their chronology and statistics. It was shown that most dynamic load tests performed in Japan have been conducted on steel pipe piles in offshore conditions for the acceptance of the tested piles. A total of 13 comparative cases of dynamic load testing and static load testing, which were conducted by JASPP, have been reviewed in detail. The following findings were presented: 1. It is difficult to predict the set-up ratio of a pile accurately. Hence, conducting a re-driving test of the pile is required to estimate the bearing capacity as well as the load-displacement curve for the pile. 2. The load-displacement curve derived fi-om the dynamic load test is in fairly good agreement with
Shaft resisitance, z (kN/m2)
-from Static load test _
I
-
Wave matching analysis with rational soil models
- - -Wave matching analysis with the smith soil mode Figure 23. Distributions oft,, derived from wave matching analyses using the S m i t h model and the rational soil models.
inside the open-ended pipe pile and the mobilization of the internal shaft resistance are taken into account, according to the modeling of the internal soil (soil plug) proposed by Randolph (1987). The soil parameters except for the maximum shaft resistance, zma,, and the maximum toe resistance, q b , were estimated from the equations listed in Table 4. The shear wave velocity, 6,of the soil was estimated Figure 24. Static load-displacement curves derived from the using the following empirical equation (Imai 1977); wave matching analyses and obtained from the static load test.
748
that obtained from the static load test until the yield load is reached, if both the tests are conducted at time points sufficiently after the end of initial pile driving.
REFERENCES Hajaslii. M . Matsuiiioto. T & Suzuki. M . 2000 Dliiaiiiic load testing on 102 steel pipe piles for bridge fouiidations on inudstone Proc 5th Int Conf 011 the .4pplicatioli of rhe Stre,r-Iim’e Theor1 to Pilec. Sao Paulo (to be published) Iiiiai. T 1977 P and S n a \ e lelocities of the ground iii Japan. Proc. 9th I C Y W E . T o b o . 1977 Matsuiiioto. T & Takei. M . 1991 Effects of soil plug on beha1 lour of drn en pipe piles. Soils and Foundations. V0l3 1. NO 2 11-31 Raiidolph. M F . 1987 Modeliiig of the soil plug respoiise duriiig pile d m iiig Proc 8th ,C E .-lsrnti Geotechtiicnl C’onf. Bangkok. Vol 2 6 1-6 1 1 Randolph. M F & Siiiioiis. H A . 1986 An iiiiproved soil inodel for one-diiiiensioiial pile drir ing anal! sis Proc 3rd 1111Coiif on \’zii?? .\Jet11 I I I Oj,?ihore Pilitig, Nantes 1-1 7 Raiidolph. M F and Decks. A J , 1992 Djnaiiiic and static soil inodels for alial pile response Proc of 3rd Iiit C ’ m f oii .4pplrcntion of Stress-11nve Tlieort to Piles. Hague 311
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This Page Intentionally Left Blank
Application of Stress-Wave Theory to Piles, Niyama & Beim (eds)02000 Balkema, Rotterdam, ISBN 90 5809 150 3
Case studies of high capacity CFA pile testing in Australia S. Baycan Pile Test International,Melbourne, Vic.,Australia
ABSTRACT: Continuous Flight Auger (CFA) piling continues to be a cost effective solution employed throughout Australia to solve many foundation engineering problems. This paper aims to describe two examples of dynamic load testing on recent CFA projects, including a comparison with a static load test, A brief description of pile construction, dynamic and static load testing for piles constructed at the new Sydney International Airport and the new Adelaide Convention Centre will be outlined.
2 GEOTECHNICAL PROFILE
1 INTRODUCTION 1.1 Sydney International Airport
2.1 Sydney International Airport
As part of the upgrading of Facilities at the Sydney International Airport in preparation for the 2000 Olympic Games, Vibro-pile Aust Pty Ltd was commissioned by Transfield Constructions to design and construct foundations for the proposed extension works. The project required constructing foundations into widely variable ground conditions overlying a weathered rock profile. Due to limitations on vibration and disturbance to underground services approximately 500 number CFA piles varying from 0.6 to 0.9m diameter were constructed for the project. The test program consisted of a series of low strain integrity, dynamic and static load testing. Geotechnical information, pile construction records and dynamic test results were used to determine a CAPWAPmodel. Model results compared well with the static load test result.
The idealised geotechnical profile within the area of piling works at Sydney International Airport was characterised by fill, overlying variable relative density sand with interbedded layers of soft clay and sands of medium density. Very stiff sandy clay was found to overly the sandstone bedrock. CPT testing was used extensively to investigate the ground conditions and a summary of ground conditions is shown in Table 1. Depth 0-4m 4-5m 5-8m
Geotechnical Profile Med. Dense - Dense SAND Soft CLAY Med. Dense SAND
1.2 Adelaide Convention Centre A combination of ‘V’ (a Vibropile proprietory displacement pile) and CFA piles were adopted by Vibro-pile Aust Pty Ltd for the foundations at the Adelaide Convention Centre project. This project required careful planning due to the structure straddling a roadway and railway lines. A total number of approximately 300 piles were constructed. An example of a dynamic test result on a 500mm diameter CFA pile and the corresponding CAPWAPmodel for such a test is described.
2.2 Adelaide Convention Centre The geotechnical profile at the Adelaide Convention Centre site consisted of a variable layer of fill up to 3m in depth, overlying very stiff silty clays to approximately 14m, overlying very dense sand and gravel. Residual siltstone - hard clays underly the 75 1
sanddgravels below 18m depth. Site investigation consisted of borehole drilling and the use of SPT testing. The shaft of pile T3 was predominantly in very stiff - hard clays with the base at 16.5m in very dense sandy gravel.
Excess 36 t Rig M R Site WHUENIION, BDElilIDE 2182 Diaieter 588 NI Total Vol 4.551 n3 Borelin 16 nin Start 87% Pile P.TEST length 16\58 n Precharge 8.284 w3 Canclin 9 Rin Stop M:56 Ilate 19/82/88
Bore
Penetrate Toque
(dmin) (wdrev) ($1 8 1 2 3 8 288 8 58
Extract Pressure Oversupply Profile (Wnin) (bar) (11 (RR) 8 2 4 6 8 1 2 3 8 48 148 258 R 258
3 PILE DESIGN AND CONSTRUCTION Careful planning and site specific constraints required consideration from the Vibro-pile designers for the Sydney Airport project. These included piling within 50m of Boeing 747 aircraft and minimising affects on underground fuel lines and other services. CPT investigations were used extensively to establish indicative pile design parameters. Individual piles were designed on a combination of shaft and end bearing. Design parameters up to 45% greater than those recommended in the geotechnical report were adopted. Elastic pile design using PIGLET was used to initially design pile groups. Pile testing was planned early on in the project to justifi seemingly ambitious design assumptions. Pile testing was critical as results were used to back-analyse pile global settlement calculations. Pile construction was digitally monitored and relayed to the office - in real time to provide added assurance of pile construction quality. Vibro-pile used a similar on-board computer system PL20 to monitor pile construction for the Adelaide Convention Centre piles. The computer system is able to display construction parameters such as depth, penetration rate, torque, injection pressure and volumetric oversupply. This data is used to obtain a better CAPWAPpile model. An example printout of the PL20 system is shown in Figure 1.
1
Figure 1 Adelaide Convention Centre PL20 pile record T3
value of 150'?Aos* (where S* is the factored design load in ultimate limit state analysis, which is usually 1.35 x working load). The test acceptance criteria consisted of: (a)
4 PILE TESTING PROGRAMME
(b)
4.1 Sydney International Airport
(c)
Every pile constructed for the new Sydney International Airport upgrade was integrity tested using low strain, frequency response and sonic echo methods. Based on integrity test data, all piles were found to be satisfactorily constructed. A total of eight piles were dynamically load tested, however for brevity the results for pile T5 only will be reviewed here. The test system for the static load test comprised of jacking against steel beams with a lOOOtonne capacity hydraulic actuator. Tension piles were used as reaction. Pile movement was monitored by string POtentiometers, with dial gauges and dumpy level readings used as backup to verify data. The loading cycle was applied in accordance with AS2159 . The load-unload and reload sequence comprised of loading in additional increments of 15% up tc 2 752
pile top settlement 4 0 m m at design working load, Net pile settlement 2 2 or Z1 < 22, equivalent to an anomaly detected with SIT). e) Impact (equivalent to the driving condition during PDA or redrive during DLT). 4 REVISION OF THE MAIN TECHNOLOGY DEVELOPMENTS He have considered gathering the information delivered by the two main sources that manufacture equipment, including hardware and software: P The European, represented by TNO Building & Construction Research, Delft, (Holland) and, P The American, represented by PDI- Pile Dynamics, Inc. Cleveland, Ohio (USA). 4.1 The main developments of TNO
The equipment produced by this R&D Institute is characterised by the generic denomination of FPDS- Foundation Pile Diagnostic System and it is described as follows (1 4):
P Equipment FPDS-0.
Likewise, the force applied on the section is given by: F = -EA du/dx F = -EA df /d(x-ct) - EA df /d(x+ct) (5) F=F +F (6) In turn, the velocity is expressed by: v = duldt = v = - c df /d(x-ct) + c df /d(c+ct) v = v +VI
(7) (8)
'These two velocities are linked to their respective f ~ r c e by s the following expressions: (9) v =Fi/Z v = - F /Z (10) where Z = EA/c Z : bar impedance. The application of the above equations to the different situations arising daily in the pile industrycoupled to the powerful new "high tech" developments- have paved the way for a rational analysis of the signals obtained from SITS, PDAs and DLT tests, as a result of which it has been able to differentiate the following cases:
Photograph 1. Equipment FPDS-0 system with an oscilloscope and Polaroid camera
P Equipment FPDS-1. In the mid 80s, the FPDS-I was developed, based on a UNIX workstation and a MC68000 (1 6 bit micro-processor chip), with built-in keyboard, disk drive, graphical display, and printedplotter. The signal was digitised (12 bit) at an early stage after retrieving the signal to ensure that all subsequent processing was of the highest possible standard. The system had certain automatic functions, but the 76@
processor) and a subsystem with an analogue-todigital (A/D) converter board and a specially designed board for integrity testing. As with FPDS-2, the subsystem could accommodate different FPDS applications, such as PDA/DLT, the Statnamic load test (STN), vibration measurements (VIBRA), and others. Each option consisted of specially designed boards, sensors, software, and accessories, an innovation begun with the FPDS-2 system. Automatic signal conditioning freed the operator from bridge balancing and scale selection. To reduce weight while conducting multiple SIT tests, the signal conditioner was built into the GRiD computer (called the Consultant System). With FPDS-3, SIT was also extended to include an optional instrumented hammer (an accelerometer fitted in a hammer head) to record the applied impact blow. The design of the hammer head, impact cap, accelerometer sensitivity, and hammer shaft stiffness produces a load pulse signal perfectly representative of the impact amplitude and phase. Since force is proportional to velocity at the pile head, the shape of the force and velocity traces are identical until a reflection is recorded. By comparing the force signal to the velocity signal, defects in the first 2 meters of the pile are revealed. The force is equivalent to the pile head impedance ’ velocity.
skilled operator could also override, magnify, or clarify a result to facilitate interpretation. The FPDS-1 system is still used today by clients for SIT.
Photograph 2. Equipment FPDS-I .
9 Equipment FPDS-2. With FPDS-2, the computer and signal conditioner were separated, a philosophy continued with subsequent systems. The FPDS-2 system, introduced in 1986, was based on a portable IBM-PC/AT compatible computer running under MS-DOS, equipped with a 20 MB hard disk for program and data storage and a floating point co-processor. For each FPDS application, specific sensors, signal conditioning subsystem, and software were available, such as Pile Driving Analysis and Dynamic Load Testing of foundation piles (PDNDLT). The FFDS-2 system did not require knowledge of electronics, requiring only a limited number of components and no oscilloscope or tape recorder. Real time presentation of signals, automatic controls, warnings, and error messages were a big improvement on earlier systems.
Photograph 4. Equipment FPDS-3 Photograph 3. Equipment FPDS-2.
9 Equipment FPDS-3. At the beginning of the 90s, the FPDS-3 system was developed, based on a GRiD computer (80386
P Equipment FPDS-4. In 1993, FPDS-4 was developed as a very robust and economically designed lunch box type system. The FPDS-4 system includes a 80486 processor, 8°-200 MB hard disk, and a built-in graphical 761
screen and key board, both sealed to protect against rain and dust.
P Equipment FPDS-5. Recently, the FPDS-5 system was developed using a notebook computer and separate signal conditioning and A/D card subsystem. Data communication from subsystem to notebook is via a PCMCIA card and cable. The FPDS-5 system is suited for PDNDLT and STN, but can also be used for SIT.
Photograph 7. Equipment FPDS-6, Field computer with SIT PCMCIA card.
P Years 1960/1979. Hardware: Development of analog instrumentation and computation technology. Software: Digital integration in the laboratory. Beta Method published.
Photograph 6. Equipment FPDS-5, Notebook with SIT PCMCIA card.
P Equipment FPDS-6. As well, TNO's latest FPDS system, the FPDS-6, represents a new approach to foundation testing equipment. The FPDS-6 is a hardware/software package for SIT, designed to work with a wide range of IBM compatible PC computers equipped with a PCMCIA card slot. The SIT hardware consists of a hammer, sensor, cable, and PCMCIA card. The PCMCIA card, measuring just 8.5 cm by 5.5 cm by 5 mm, incorporates all data acquisition (300 kHz, 14 bit) and signal conditioning electronics. Besides the enormous space and weight savings, this innovation dramatically increases overall system reliability and service life. The user can use his own notebook computer or a hand-held Hardbody PC. All input is via an easy to use touch screen. Knowledge of electronics is unnecessary-all electronic functions are controlled by the computer. Advice and warnings are generated by the system. Signals and data are stored automatically and can be easily recalled. Automatic reporting capabilities greatly reduce reporting time.
P Years 1983. Hardware: a) Data acquisition: For low strain, highly sensitive accelerometer and power supply; digital data acquisition and integration with Pile Driving Analyzer (PDA); digital display on oscilloscope. b) Kecording: Oscilloscope, photos of oscilloscope traces, analog FM magnetic tape, X-Y plots. Software: Analysis in time domain
k Year 1985. Hardware: IBM PC or compatible 8087 math coprocessor, 5 12 kRam, A/D board (acceleration records integrated and digitised) , signal amplifier. Software: m e g r i t y Program, exponential amplification, averaging of records, smoothing.
P Year 1989. Hardware: PIT test equipment: PC + signal conditioning; PIT processor and plotter or graphics printer in field; permanent storage on disk. Software: Analysis in frequency domain.
4.2 The main developments of PDI
P Year 1990. Hardware: PIT-SC: battery powered, signal conditioning and printer built in a 80286-based PC computer. Instrumented hammer.
The equipment manufactured by this company and its respective innovations are described in the following list (1 5): 762
5.1.2 Problems associated to drived piles
Software: Built in time and frequency domain software. Impedance profile. PITWAP (calculates pile shape given a soil model, based on wave equation).
5.1.3. Problems associated to external causes in the manufacturing process.
>
A detailed description of the problems and the methods available for their detection has been clearly exposed by M.J. Turner (19) and by J.J. Goldemberg & H. Goldemberg (20). As typical examples of the anomalies detected in foundations built with cast-in-place piles the following deserve to be highlighted:
Year 1992. Hardware: PIT Collector, touch screen, 16 bit A/D converter, high and low pass data filtering; signal amplification; interface to plotter/printer; full day battery; storage of 350 piles. Software: Built in analysis software in time domain. Post processing in frequency domain.
>
>
Year 1996. Hardware: PIT Collector memory expanded; improved display capabilities. Software: PITSTOP (frequency domain, 2 velocity, multiple plots, impedance profile).
Case History I.
Corresponds to a sector belonging to a group of Towers of 1 1 stories destined to economic housing, built in the south zone of Buenos Aires (Argentina) and where, due to contractual reasons, two of the mentioned buildings remained for a long time with their foundations (piles, caps and foundation beams) partially built without their superstructure completed. The geotechnical profile of the building site (see Figure 2) shows a first layer of low plasticity clays (CL) up to 4.00/4.50m depth with N value (SPT) between 3 and 16, laying on a non plastic silty strata (ML) with similar respond to the SPT test, up to a depth of 6.00m. From here onwards, the soils prcsented a strong preconsolidated characteristic with N>40, up to well below 10.00m from zero level. At the moment of the investigations, the water table level was 2.00/32.50m below ground level. The foundation system used consisted of displacement piles by driving a recoverable close ended caisson and casting the concrete while the mould was pulled out with the help of a vibratory hammers. Their nominal diameters were: 300/350 and 400mm and length of 4 to 6m. When the continuation of the building was decided, by the Main Building Contractor’s Structural Advisor, the Quality Control upon the total number of the piles
P Year2000. Hardware: Improved memory (1000 piles) and processing speed. Software: PIT-W (Windows) 5 COMMON PATOLOGIES IN DEEP FOUNDATIONS. In order to describe the principle causes of pathology detected in the foundation structures, it is convenient to list them into: 5. I Problems originated in its construction.
5.2 Problems derivedjkom its use during lif. time of the structure.
5.3 Problems due to geotechnical- environmental phenomenon. In this paper a reference to the problems generated during the construction of cast-in-place piles will be made for being the most common in our deep foundation market and the ones which have more possibilities of showing anomalies, according with the authors experience. In general the problems that can turn out of the construction of different piles and deep foundations, had been described in a number of publications: texts, seminars, conferences, etc. (16), (17), (18). In order to identify those causes of damage, problems can be classified in the following way: 5.1. I Problems piles.
associated
to
cast-in-place
Figure 2. Existent geotechnical profile in the building site.
763
of both buildings through the performance of the Sonic Integrity Tests was arranged. In accordance with it, the 108 piles of each building were tested, upon preparing its in order to eliminate all the material “that could be loose”, concrete contaminated by the soils, cracks caused by the trimming work due to use of inadequate techniques, etc. That could originate false signals. Given that some of the piles had the cap already built, they were tested under two different conditions: a) from the upper face of the cap and b) from a concrete brick expressly attached to the upper sector of the shaft - beneath the lower face of the c a p by means of special epoxy resin of quick curing due to the groundwater proximity. Given that, the equipment FPDS-3 was used as well as its corresponding software developed by TNO - Building & Construction Research; the records obtained were adequately “filtered” to take into consideration the presence of the caps, in the indicated cases. The preliminary analysis of the signals revealed a real alarming picture, as they previously showed a number of piles with anomalies which easily exceeded- the usual expected percentages for this type of piles. In the view of the possibility that the diagnosis were influenced by a mistaken interpretation of the geotechnical data provided by the Customer, the performance of new reliable Geotechnical Studies were arranged, as to being able to apply- in a second stage- the Signal Matching Technique (21). Likewise, and giving the worrying situation (because it could lay open to discussion the diagnosis by the use of technology that was rather unknown in our environment at that time), the following measures were adopted: a) Execution of a new Geotechnical Study, as it is described above. b) Take new signals in the total number of piles. c) Apply a Signal-Matching Technique on some of the observed piles. d) Ask for a second independent opinion to TNO- Building & Construction Research, in accordance with the agreement we kept with the mentioned Institute. As a consequence of the previous, we arrived to the following conclusions: i.- The original geotechnical study contained no mistakes that would have led to erroneous conclusions when modelling the soil to apply the TNOWAVE program option Signal Match. ii.- The repetition of the Sonic Integrity Tests showed that the preliminary diagnosis, derived from the first series of signals, kept its force; it is summarised as Table 2 shows. To illustrate the example, some typical signals corresponding to “piles with anomalies” (Figure 3) and “normal pile” (Figure 4) corresponding to Tower A are enclosed
Table 2. Summary of tests on piles. TOWER PILES No Total w/Anomalies A 108+6* 72+6* 26 B 108 Note: * Piles replaced by order of the Main Building Contractor before SIT testing.
5 7m
A
.
f 57rr
0. Icm/
5’6m
0 1 2 3
4
5
Pile 65 29 Jan 94
6
7
8
9
:
v2 a
G
f 5
4000P/s
I
I
sr
e x p 70
Figure 3. Pile #68; anomaly close to the 3m.
4 7m
I a
P i l e 12 26 Jan 94
1
2
3
4
5
6
7
a
9
4000m/s
__
1
v2 0
0 f.
a
sr
exp. 5
Figure 4. Pile #12; without anomalies.
iii.- From the number of Simulation Analysis done, on the behalf of the TNO specialists as well as the author and his colleagues- in which the signal matching techniques were applied through the TNOWAVE program option SIT-SM -,we arrived to the conclusion that most of the “suspicious” piles presented anomalies that were incompatible with their security (stability and life time). The respective graphics are included (Figures 5 and 6). iv.- Based on the conclusions that were obtained, we came to agreement with the main building contractor: to carry out the excavation and the drain of the staircase and elevator sector of Tower A in order to be able to observe the condition of the piles involved. In spite of the difficult work, they were performed efficiently, allowing the personal observation- on the behalf of all the people involved in the construction- of the anomalies which their prediction, based on the Sonic Integrity Tests and the 764
Figure 5 . Signal Match of pile #68 (anomalous).
Figure 6. Signal Match of pile # I 2 (normal).
adequate application of the Signal Matching Techniques, allowed the early detection of serious problems derived from the lack of geometricstructural integrity of a large number of piles in both buildings (see photographs 8 and 9). v.- In order to solve the structural problems, the replacement of the defective piles through the installation of additional ones using the following construction methodology was determined: a) Installation of a temporary open ended caisson by means of a vibratory hammer up to the indicated depth by the complementary studies. b) Inside cleaning. c) Installation of the reinforcements. d) Casting concrete with tremie pipe. e) Retrieving the caisson with vibratory hammer. f ) Control of the substitute piles through the Sonic Integrity Tests (SIT). vi.- The substitute piles were exposed to the Quality Control Tests by means of the indicated technique, and the obtained results are showed in the table:
Photograph 9. Characteristics of the damages.
Table 3. Summary of the additional piles.
TOWER
N" A B
k Case History I1
PILES Total 49 23
w/Anomalies 0 0
Corresponds to a group of three luxurious buildings of thirty stories high built in the north of Buenos 765
Aires city (Argentina); the Architecture Project foresaw the building of a basement that held the totality of the place (a block). Taking into account the bad quality of the upper soil layers (see Figure 7 ) it was decided that the foundation should be with displacement cast-in-situ piles with extended bulb (Franki type). For building reasons, the piles were installed from ground level and casted up to the basement level, which was located beneath the level of general excavation and close to the groundwater level. Because of budget reasons, excavations for the caps and foundation beams construction were done by reduced sectors- without a general groundwater abatement.
4.61~
0.I c d 4.681
0.lcn/s 0
1
2
3
4
5
6
7
8
P i l e 581
4000a/s
21 dug 94
exp: 20
9
1
v2.0
0 f: I
sr
Figure 8. Pile #55:anomaly at 1.5m
for the already mentioned reasons, was taken as model pile (Figure 9).
3’7m
O.lCrn/
0
1
2
3
4
5
6
7
8
P I ] @ 38
4000m/s
17 ADP 94
exa: 2
9
1
I
v2.0
0
I:
sr
Figure 9. Pile #38: without anomalies. Figure 7. Existent Geotechnical Profile in the building site
That brought the impossibility of an appropriate coordination that allowed carrying out and efficient quality control of the totality of the piles because of which, the Engineer and the Construction Company determined that the work was done in a partial and random way as the progress of the work plan of the building will allow, thing that gave place to a statistical pseudo-control. In accordance to the previously mentioned, a small amount of Sonic Integrity Tests were done through the usage of the already mentioned FPDS-3 equipment. Nevertheless the limited quantity of available signals, made the adoption of the statistical standard signal corresponding to the family of records representative of the construction (or the sector) problems, anomalies in several of the registered signals were detected. The main one is showed next (Figure 8). The Simulation program TNOWAVE option SIT-SM (Signal Matching Technique) was applied on that signal and the pile NO38 that is located near,
766
In accordance with the usual confrontation of opinions (the Engineer-Main Building ContractorFoundations Subcontractor-Geotechnical Consultant-Structural Advisor-etc.) when a customer is informed that in his site a pile with anomalies has been detected. The conclusion obtained was similar to the one obtained previously, as Figures 10 and 11 show.
Figure 10. Signal Match of pile #55 (anomalous).
Not-withstanding the previous, and in the view of the disbelief showed by the Engineer, we proposed the excavation of the pile and it could be seen a series of anomalies (deep craks, loss of coating, etc.) which are indicated in Photograph 10.
quality of the soil in the building site (see Figure 12) the planners decided that the foundation should be piled. Franki type piles were adopted.
Figure 1 1. Signal Match of pile #38 (normal).
Figure 12. Geotechnical profile
Figure 13. Pile #1: anomaly at 5.501~1
+--@-I
+ Photograph 10. Pile #55. View of the observed damages.
er
P Case History 111 This construction site, which took place in the north zone of the Gran Buenos Aires (Argentina), constituted by a number of 10 storey towers destined to economic housing. Taking into account the bad
Figure 14.
767
ano'nalyat 8.10m.
-
The total number of the piles of the construction (420) was tested as the Foundation Company was completing the works corresponding to each of the towers. They had previously done the trimming work and cleaning after seven days of casted, in order to obtain an adequate resistance and to be apt to transmit the stress wave (SW) generated by the impact of the hammer. Only two piles, No 1 of the Tower C2 and the No 154 corresponding to Tower A, thrown anomalous signals so a more detailed analysis through the application of the Signal Matching Technique had to be done. Due to the proximity of the observed piles, Pile 44 from the Tower C2 was adopted as "model pile"; the respective signals are showed next: (see Figures 13 to 15). As a consequence of the application of the TNOWAVE program SIT-SM option, the graphics that shows figures 16 to 20 were obtained. The detailed analysis allowed us arriving to the following conclusions: a) Pile 1 (Figs. 13 and 16). This is a complicated signal because the original pile, as a consequence of having problems with its installation, had to be replaced by redriving the caisson in the same place with the concrete of the first one still fresh; as a consequence, the pile shape turned to be highly irregular showing increases and reductions in its diameter. The signal match determined a variable increase in the pile diameter- from 350mm to 450mm in relation to the depths between 2.5 and 5.2mfollowed by an abrupt reduction to the nominal diameter starting at 5.6m. The corresponding reflection of the pile top is clearly marked at an approximately depth of 13.0, while the signal that is detected at a depth of 11.5m corresponds to the repetition of the previous signal of increase and the subsequent section reduction. b) Pile No 154 (Figs. 14 and 17). The SignalMatch (SM) established the presence of a "larger anomaly" to an approximately 8.1m corresponding to a discontinuity, reason why the stress wave that was generated could not continue its way further than that depth, giving place to the "reflection corresponding to the end"; the rest of the diagram corresponds to the first repetition of the signal. From the Sonic point of view and consequently from its capability to transmit shear stresses, is as if the pile were 8.lm long even how, probably, it was driven to its nominal depth, but that discontinuity could have been a crack or an important soil inclusion as it happened in the Case History I. The problem was solved when the Engineer arranged the replacement of the pile by other two of smaller dimensions and the modification of the pile cap and foundation beams.
I
Figure 15. Pile #44: without anomalies.
-eoo1
.
,
0 0
-----
20
CDlLUldlCB
1 0
an0
- - -
,
,
6 0
6 0
I
100
,
,
,
120
260
160
I
I
I80
200
sec I 10-3 .~
m?aI"PE(I "ClOCillCS
Figure 16. Signal Match of pile # I .
-1 2
t
-1
-2.01 00
,
,
20
.___. C*,C"llrfCd
4 0
an0
- .-
,
,
,
69
8 0
100
12.0
,
,
140
160
P P 0 ~ " P C ~"CIOCIIIES
J ZWO
IS0 5rC
I
10-3
Figure 17. Signal Match of pile #154.
-2 0 0 0
2 0
4 0
8.0
0 0
10.0
I20
140
160
100
a 0
Figure 18. Signal Match of pile #44
768
6 CONCLUSIONS.
neering. Design and practice guide”. (Thomas Telford, London, 1996). (9).- Institution of Civil Engineers: “The Value of Geotechnics in Construction”. (Construction Research Communications, London, 1998). (10) Love, A.E.H.,: “A Treatise on the Mathematical Theory of Elasticity”. (4th. Edition, New York, Dover Publications, 1944). (1 1) Timoshenko, S. & Goodier, J.N.: “Theory of Elasticity”. (2nd. Edition, McGraw-Hill Book Company, Inc. 195 1). (12) Timoshenko, S. & Young, D.H.: “Vibration Problems in Engineering”. (3rd. Edition, D. Van Nostrand Conpany, Inc. 1955). (13) Voitus van Hamme, G.E.J.S.L.; Jansz, J.W.; Bomer, H. & Arentsen, D.: “Hydroblock and improved piledriving anilisis”. (De Ingenieur, no 18, ~01.86,1974). 14) Middendorp, P. Private communication, June 2000. (1 5 ) Beim, G. K., Private communication, June 2000. (1 6) Thorburn, S. & Thorburn, J.Q.: “Review of the problems associated with the construction of cast-in-place concrete piles”. (DOE and CIRIA Piling Development Group, Report PC2, London, 1977). ( I 7) Healy, P.R. & Weltman, A.J.: “Survey of problems associated with the installation of displacement piles”. (DOE and CIRIA Piling Development Group, Report PG8, London, 1980). (18) Fleming, W.G.K.; Weltman, A.J.; Randolph, M.F. & EIson, W.K.: “Piling Engineering”. (Surrey University PresdBlackie and Son Ltd., London, 1985). (1 9) Turner, M.J.: “Integrity testing in piling practice” (CIRIA Report 144, London, 1997). (20) Coldcmberg, J.J. & Goldemberg, H.; “Patologia en fundaciones profundas. Origen y clasificacion”. (Memorias del V Congreso Iberoamericano de Patologia de las Constnicciones, VlI Congreso de Control de Calidad CONPAT 99, Montevideo, Uruguay, 1999). 6) (21) Goldemberg, J.J. & Goldemberg, H.; “Confiabilidad de 10s ensayos de integridad sobre pilotes por mCtodo sonico - SIT”. (Memorias del I Congreso Paraguay0 de Ingenieria Geotecnica, IV Jornadas Gcotkcnicas Estnicturales, 1 ra. Reunion de Ingenieria Geotecnica del MERCOSUR - 1er. COPAINGE, Asuncion, Paraguay, 1997).
The results obtained in the Quality Control done since the beginning of the application of this technology, in Argentina as in the rest of the world, allow establishing- without doubt- that the Sonic Integrity Tests are an essential tool to know the condition of a pile, new ones and old ones (case of pre-existent old foundations recycling or in cases of Forensic Engineering) As in every indirect investigation technique, where the diagnosis is obtained through analyzing a measurements with sensitive instruments or image analysis of signals corresponding to physicmechanic phenomenon, the unavoidable requirements are: > The availability of appropriate equipment, with proved and reliable hardware and software. P Skilled operators P Periodic calibration of sensors. P An exhaustive and reliable Soil Investigation. P An adequate control of the execution of the foundations. P Full knowledge of the Foundation Engineering, Geotechnical Engineering and Foundation Dynamics by the Engineer in charge of the Integrity Tests. P Independent criteria and ethical integrity on its behalf.
7 BIBLIOGRAPHIC REFERENCES. ( I ) Bredenberg, H. (Editor); “Application of Strcss-Wave Theory on Piles” (Proceedings of the International Seminar on the Application of Stress-Wave Theory on Piles, Stockholm, Sweden, 1980). (2) Holm, G.; Bredenberg, H. & Gravare, C.J. (Editors); “Application of Stress-Wave Theory on Piles” (Proceedings of the Second International Conference on the Application of Stress-Wave Theory on Piles, Stockholm, Sweden, 1984). (3) Fellenius, B.H. (Editor); “Application of Stress-Wave Theory to Piles” (Proceedings of the Third International Conference on the Application of Stress-Wave Theory to Piles, Ottawa, Canada, 1988). (4) Barends, F.B.J. (Editor); “Application of Stress-Wave Theory to Piles” (Proceedings of the Fourth International Conference on the Application of Stress-Wave Theory to Piles, The Hague, The Netherlands, 1992). 5 ) Townsend, F.C.; Hussein, M.H. & McVay, M.C. (Editors); “Stresswave ‘96” (Proceedings of the Fifth International Conference on the Application of Stress-Wave Theory to Piles, Orlando, Florida, USA, 1996). (6) PEDIR INFORMACION A SUSUMU. Stresswave ‘2000” (Proceedings of the Sixth International Conference on the Application of Stress-Wave Theory to Piles, San Pablo, Brasil, 2000). (7) - van Koten, H. & Middendorp, P.: “Testing of Foundation Piles” (Delft University of Technology, Delft, The Netherlands, HERON, vol. 26, No 4, 198 1). (X).- Institution of Civil Engineers: “Creating value in Engi-
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Application of Stress-Wave Theory to Piles, Niyama 8 Beim (eds)02000 Balkema, Rotterdam, lSBN 90 5809 150 3
Author index
Abreu, L. 201 Amir, J.M. 313 Andreo, C.S. 261,389 Antoniutti Neto, L. 707 Aoki, N.241,375,457,635 Axelsson, G. 665 Ayasrah, I. M. 233 Aye,Z.Z. 163, 171 Balech, J. 241 Barbosa, 6 . E. 429 Barends, E B. J. 541 Baycan, S. 75 1 Beim, J.W. 127 Benamar, A. 1 17,255 Bernardes, G.P. 261,389,707 Bettess, F? 495 Boonyatee, T. 563,569 Bnino, D. 47 Cai, L.B. 11 Camapum de Carvalho, J. 157 Cannon, J.G. 393,399 Chambers, W.G. 407 Chamecki, E! R. 707 Chen, B. 65 1 Chen, R.P. 29,517 Chen, Y. M. 29,5 17 Chernauskas, L.R. 223 Cho, C.W. 41,47 Cintra, J.C.A. 375,457 Costa, C. M.C. 429,441 Courage, W. M.G. 575 Cunha, R. I? 157 Danziger, B. R. 657 de Albuquerque, F? J. R. 677 de Campos, G.C. 429,435 de Carvalho, D. 677 DiMillio, A. E 223
Dyminsky, A.S. 127
Justason, M. D. 609
Ealy, C.D. 223 Esposito, G. 575
Kalinowski, M. 267 Karkee, M. B. 689 Kawabata, N. 335,583 Kikuchi, Y. 599 Kimura, M. 563,569 Kirsch, E 249 Kita, N. 41 1 Klingberg, D.J. 403,407,715 Korkeakoski, F? 415 Kormann, A.C. M. 707 Kusakabe, 0.335,591
Fellenius, B. H. 3 13 Ferreira, J.S. 657 FOB,S.B. 157 Fujita, K.335,451,469 Goble, G.G. 3,305,327 Goldemberg, H. 345,7 19 Goldemberg, J.J. 345,719,757 Gonqalves, C. 261,389 Gutikrrez, A. 201 Hajduk, E.L. 541 Hannen, W.R. 503 Hart, L.J. 223 Hartikainen, J. 415 Hasard, D. 201 Hayashi, M. 583,697,741 Hilmi Acar, M. 51 1 Hintze, S. 665 Hoffmann, Ch. 201 Holeyman, A.E. 479,725,733 Holscher, P. 541 Horiguchi, T. 689 Huch, T. 249 Husein Malkawi, A. I. 233 Hussein, M. 91 Huybrechts, N. 725,733 Imada, K. 179 Iskandarani, W. M. 91 Janes, M.C. 609 Joer, H.A. 47 Jokiniemi, H. 415 Jonker, G. 135 77 1
Liang, R.Y. 121 Lee, J.-S. 421 Lee, M.W. 41 Lee, S.-B. 99 Lee, W.-J. 99 Lee, Y. -N. 42 1 Legrand, C. 725,733 Liang, L. 53 Liang, R.Y. 461 Likins, G.E. 205,211,327 Lima, E M.A. 375 Liu Xi-An 683 Lucieer. W. J. 65 Mackenzie, P. 403,7 15 Maertens, J. 725,733 Matsuda, Y. 187 Matsumoto, T. 179, 187,335,583, 59 1,697,741 Maung, A.W. 163,171 Michi, Y. 187 Middendorp, E! 55 1,609,617,625 Morgano, C. M. 205 Mukaddam, M.A. 9 I Mullins, A.G. 609
Nakata, Y. 179 Navajas, S. 429,435 Navaneethan, T. 171 Nawari, N.O. 121 Nishimura, S. 335,411,591, 599,741 Nishiumi, K. 591,741 Niyama, S. 429,435 Oh, J.-H. 99 Ohno, M. 469 Okahara, M. 335
Rausche, E 53,59,75,205, 21 1,327 Restrepo, C. 219 Robinson, B. 53,59 Rodatz, W. 249 Romanel, C. 127 Romell, J. 145 Roth, B.C. 503 Russo Neto, L. 707
Padcowsky, S.G. 223,28 1,541 Paraiso, S.C. 429,441 Parente-Ribeiro, E. 127 Pereira, J. H. E 157 Pinto Soares, E. 441 Pinto, EL. 275 Piscsalko, G. 205 PlaBmann, B. 249
Salumoto, J. 41 1 Seidel, J.P. 59, 193,267, 319 Selby, A. R. 495 Shibata, A. 583,741 Skov, R. 107 Stenersen, K.L. 281 Stevens, R.E 17,351 Sugimura, Y. 689 Suzuki, M. 697 Svmkin, M.R. 35,107,113, 503,525
Ramshaw, C. L. 495 Randolph, M.E 41,47
Takeda, T. 41 1 Tanseng, I? 163
772
Tatsuta, M. 599 Thasnanipan, N.163, 171 van Foeken, R. J. 135,575,625 van Ginneken, G.J. J. 6 17,625 Viking, K. 533 Wakiya, Y. 583,741 Wu, J. 673 Wu, S. 383 Xi Liang 369 Xiao, L. 383 Xu, D. 383 Yang Wu 153 Yoshizawa, Y. 583,591 Zhang, F. 563 Zhang, Y.-N. 683 Zheng, J. M. 65 1 Zheng, Y. M. 65 1 Zhou, G. 673 Zhu, B. 5 17
