SAVEZ GRAEVINSKIH INENJERA SRBIJE ASSOCIATION OF CIVIL ENGINEERS OF SERBIA

INENJERSKA KOMORA SRBIJE SERBIAN CHAMBER OF ENGINEERS

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IZDAVA / (PUBLISHER): Savez graevinskih inenjera Srbije / Association of Civil Engineers of Serbia Beograd, Kneza Miloa 9/I, Tel/Faks: (011) 3241 656 PROGRAMSKI ODBOR / (PROGRAMME COMMITTEE): KOPREDSEDNICI / (CO-CHAIRS): Prof.dr Radomir FOLI, Novi Sad, Srbija Prof.dr Milan MAKSIMOVI, Beograd, Srbija LANOVI / (MEMBERS): Prof. Dr. Heinz BRANDL, Wien, Austrija Prof. Dr. Luvig TRAUNER, Maribor, Slovenija Prof. Dr. Lidija ZDRAVKOVI, London, U.K. Prof. Dr. Tanja ROJE-BONACCI, Split, Hrvatska Prof. Dr. Ivan VRKLJAN, Rijeka, Hrvatska Prof. Dr. Vasil VITANOV, Skoplje, Makedonija Prof. Dr. Adnan IBRAHIMOVI, Tuzla, BiH Prof. Dr. Asterios LIOLIOS, Xanthi, Grka Doc. Dr. Georgi FRANGLOV, Sofia, Bugarska Prof. Dr. Zvonko TOMANOVI, Podgorica, Crna Gora Prof. Dr. Slobodan ORI, Beograd, Srbija Dr. Nenad UI, Beograd, Srbija Prof. Dr. Milinko VASI, Novi Sad, Srbija Prof. Dr. Mitar OGO, Novi Sad, Srbija Prof. Dr. Petar SANTRA, Subotica, Srbija Miroljub SAMARDAKOVI, Ni, Srbija EDITOR / (Editor in Chief): Prof.dr Radomir FOLI TEHNIKI UREDNIK / (Editor): Aleksandar UKI Svi radovi u ovom zborniku radova su recenzirani. Stavovi izneti u ovoj publikaciji ne odraavaju nuno i stavove izdavaa, naunog komiteta ili editora. TIRA (Circulation): 200 TAMPA: Akademska izdanja, Zemun

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SAVEZ GRAEVINSKIH INENJERA SRBIJE

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SRPSKO DRUTVO ZA MEHANIKU TLA I GEOTEHNIKO IENJERSTVO

ZBORNIK RADOVA

PETO NAUNO-STRUNO MEUNARODNO SAVETOVANJE

GEOTEHNIKI ASPEKTI GRAEVINARSTVA

FIFTH INTERNATIONAL CONFERENCE GEOTECHNICS IN CIVIL ENGINEERING

CONFERENCE PROCEEDINGS

Editor: Prof. dr Radomir Foli

Sokobanja, 29. - 31. oktobar 2013.

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ORGANIZATORI SAVETOVANJA / (CONFERENCE ORGANISERS): Savez graevinskih inenjera Srbije (Beograd), Srpsko drutvo za mehaniku tla i geotehniko inenjerstvo (Beograd) ORGANIZACIONI ODBOR / (ORGANISING COMMITTEE): PREDSEDNIK (Chairman): Prof.dr Radomir FOLI, dipl.in.gra., Novi Sad SEKRETAR (Secretary): Nevena VUJADINOVI, Beograd LANOVI (Members): Dr Vencislav GRABULOV, dipl.in.tehn., Beograd, Srbija Milutin IGNJATOVI, dipl.in, Beograd, Srbija Mr Dragan ZLATKOV, dipl.in.gra., Ni, Srbija Jovo SMILJANI, dipl.in.gra., Novi Sad, Srbija Milica TRIFKOVI, dipl.in.gra., Beograd, Srbija Dubravka PETKOVI, dipl.in., aak, Srbija ODRAVANJE SAVETOVANJA SU POMOGLI / (SPONSORED BY):

Ministarstvo prosvete, nauke i tehnolokog razvoja Republike Srbije Inenjerska komora Srbije Institut IMS a.d. Beograd Lipex d.o.o. Beograd Geoestetika d.o.o. Beograd Geoput d.o.o. Beograd Projektinenjering Tim d.o.o. Ni Sigmainenjering d.o.o. Novi Sad Benevento Ums d.o.o. aak Saobraajni institut CIP d.o.o. Beograd

Slika na koricama: iskop za izgradnju garae "Pionirski park" u Beogradu (autor prof. dr Milan Maksimovi)

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S A D R A J C O N T E N T S

Radovi po pozivu / Keynote Papers 1. H.Brandl (Vienna Austrija)

BOX SHAPED DEEP FOUNDATIONS TO IMPROVE THE BEARING-SETTLEMENT BEHAVIOUR OF STRUCTURES ......................................................................1

2. L.Trauner, A.trukelj, M.Punder, B.Macuh (Ljubljana Slovenia) A STATIC LOADING PILE TEST AT THE SERVICEABILITY LIMIT STATE ........................................25

3. T.Roje-Bonacci (Split Hrvatska) PRIRODNE BRANE S OSVRTOM NA NAJVEU POZNATU BRANU USOI U TAJIKISTANU ..............................................................................................................................................43

4. I.Vrkljan (Rijeka Hrvatska) MEHANIKA STIJENA 50 GODINA NAKON OSNIVANJA ISRM-A ......................................................53

5. E.Mandi (Tuzla BiH) SLIJEGANJE TERENA GRADA TUZLA .......................................................................................................61

Tematska oblast 1 / Topic 1 NORMATIVI TEHNIKI PROPISI U GRAEVINSKOJ GEOTEHNICI U SVETLU USAGLAAVANJA SA EN STANDARDIMA GEOTECHNICAL STANDARDS AND REGULATIONS

6. J.Papi, V.Prolovi, Lj.Dimitrievski (Skoplje Makedonija, Ni - Srbija)

PREDLOG ZA PRORAUN POTPORNIH ZIDOVA U REGIONU PREMA EVROKODU 7 ..................................................................................................................................................75

Tematska oblast 2 / Topic 2 STANJE GEOTEHNIKE U NAOJ ZEMLJI STATE OF GEOTECHNICAL ENGINEERING IN SERBIA

7. V.Vujani, M.Joti (Beograd Srbija)

RAZVOJ GEOTEHNIKE U PUTARSTVU (1955 2013) ..............................................................................83

Tematska oblast 3 / Topic 3 GEOTEHNIKA U PROJEKTOVANJU I IZVOENJU OBJEKATA INFRASTRUKTURE GEOTECHNICAL ASPECTS IN INFRASTRUCTURE

8. M.Memi, R.Foli, A.Ibrahimovi (Lukavac BiH, Novi Sad Srbija, Tuzla BiH)

UTICAJ PROMJENE PARAMETARA TLA NA POMIJERANJE ARMIRANOBETONSKIH DIJAFRAGMI ......................................................................................................93

9. V.Anelkovi, D.Divac, .Lazarevi, V.Nedovi (Beograd Srbija) ISPITIVANJE KARAKTERSITIKA SMICANJA NA KONTAKTU BETON-STENSKA MASA ............................................................................................................................103

10. Z.Kovrlija, A.Tomanovi (Beograd Srbija) FUNDIRANJE MOSTA "M 13" PREKO KIJEVSKOG POTOKA, SEKTOR B5.1, OBILAZNICA OKO BEOGRADA ................................................................................................................113

11. V.Bogdanovi (Beograd Srbija) GEOTEHNIKE PODLOGE ZA GLAVNI PROJEKAT CEVOVODA 300 mm U VRANIU ...................................................................................................................................................119

12. P.Petronijevi, V.Prolovi, S.Zdravkovi (Ni Serbia) POOR FOUNDATIONS AS A CAUSE OF COLLAPSE OF THE LATTICE MAST ON VRTOP PEAK ..............................................................................................................................125

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13. M.Vasi, M.ogo (Novi Sad Serbia) GEOTECHNICAL CONDITIONS FOR THE CONSTRUCTION OF A NEW DRINKING WATER PLANT IN ZRENJANIN ............................................................................................133

14. N.uri (Bijeljina Republika Srpska BiH) GEOTEHNIKA ISTAIVANJA TERENA NA LOKACIJI POSTROJENJA ZA PREIAVANJE OTPADNIH VODA V.OBARSKA KOD BIJELJINE...................................................139

15. Z.Tali, .erimagi (Sarajevo BiH) GEOTEHNIKE KARAKTERISTIKE TERENA I PRORAUN DOZVOLJENE NOSIVOSTI NA LOKACIJI MOSTA BR. 1, AUTOPUT KORIDOR Vc, DIONICA TARIN KONJIC, PODDIONICA TARIN ZUKII ............................................................................147

16. Z.Tali, .erimagi (Sarajevo BiH) GEOTEHNIKE KARAKTERISTIKE TERENA I PRORAUN DOZVOLJENE NOSIVOSTI NA LOKACIJI MOSTA BR. 2, AUTOPUT KORIDOR Vc, DIONICA TARIN KONJIC, PODDIONICA TARIN ZUKII ............................................................................157

17. D.Peco, I.Bojovi, S.ijan, M.Savi, V.Jovanovi, N.Lazi (Beograd Srbija) FUNDATION OF RAIL WAY - ROAD BRIDGE OVER THE DANUBE IN NOVI SAD .........................163

Tematska oblast 4 / Topic 4 GEOTEHNIKI ASPEKTI GRAENJA U URBANIM SREDINAMA GEOTECHNICAL ASPECTS OF CONSTRUCTION IN URBAN AREAS

18. .ugi, M.Raki (Beograd, Loznica Srbija) PROJEKAT SANACIJE TEMELJA ZGRADE PRIRODNO-MATEMATIKOG FAKULTETA NEKI ASPEKTI IZVOENJA ............................................................................................171

19. D.Zlatkov, P.Petronijevi, V.Prolovi (Ni Serbia) FOUNDING OF THE SHALLOW ARCH IN POOR GEOTECHNICAL CONDITIONS ...........................179

20. R.orevi (Beograd Srbija)TEMELJENJE UZ PROSTORNA OGRANIENJA ....................................185

Tematska oblast 5 / Topic 5 ISTRANI RADOVI, KARAKTERISTIKE TLA I STENA, KARAKTERIZACIJA I KLASIFIKACIJA TERENA SITE INVESTIGATIONS, CHARACTERIZATION OF SOIL AND ROCK

21. M.Prica, K.okovi, N.ui, D.Berisavljevi (Belgrade Serbia)

IN SITU TESTING OF SOILS BY SCREW PLATE LOAD TEST (SPLT) .................................................191 22. K.okovi, L.aki, N.ui (Belgrade Serbia)

ASSESSING SOIL DISPERSIVITY BASED ON CLASSIFICATION TESTS ...........................................197 23. S.Samardakovi, M.Samardakovi, R.Foli (Ni, Novi Sad Srbija)

DINAMIKA PENETRACIONA ISPITIVANJA I MEUSOBNE KORELACIJE ....................................205

24. S.Krsti, M.Ljubojev, V.Ljubojev, D.Tai (Bor Srbija) GEOTEHNIKA ISTRAIVANJA TERENA NA TRASI IZMETANJA KOLEKTORA - FLOTACIJSKO JALOVITE VELIKI KRIVELJ ..............................................................213

25. E.Mandi, K.Mandi, E.Babaji, A. Ibrahimovi, E.Mandi (Tuzla BiH) KARAKTERISTIKE PERIDOTITA SERPENTINITA KOD IZGRADNJE GEOTEHNIKIH OBJEKATA ......................................................................................................................219

Tematska oblast 6 / Topic 6 MODELI GEOMATERIJALA I NUMERIKE METODE GEOTECHNICAL MATERIAL MODELS AND NUMERICAL METHODS

26. G.Hadi-Nikovi, S.ori (Belgrade Serbia)

ULTIMATE BERING CAPACITY IN UNSATURATED SOILS .............................................................225

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27. B.Foli, M.osi, .Lainovi (Beograd, Loznica - Novi Sad) NDA VI MOSTA FUNDIRANOG NA IPOVIMA PREKO PE-IPSILON KRIVIH ZA PESAK PREMA RISU ..............................................................................................................233

28. L.Zdravkovi (London UK) THE USE OF THE MOHR-COULOMB MODEL IN GEOTECHNICAL ENGINEERING PRACTICE ..........................................................................................................................243

Tematska oblast 7 / Topic 7 PREDVIANJE I REZULTATI OSMATRANJA OBJEKATA OPSERVACIONI METOD OBSERVATIONAL METHOD, PREDICTION AND MONITORING

29. E.izmi, A.Skeji, D.Ljubuni, A.Bali (Sarajevo BiH) MONITORING I NUMERIKO MODELIRANJE VREMENSKI OVISNIH POMJERANJA U KLIZITU .........................................................................................................................249

Tematska oblast 8 / Topic 8 POBOLJANJE TLA, ARMIRANJE, INJEKTIRANJE, DRENAE I DRUGO SOIL AND ROCK IMPROVEMENT

30. M. Vukievi, S. Mara-Dragojevi, S. Jockovi, M. Marjanovi, V. Pujevi (Beograd Srbija) STABILIZACIJA ALEVRITA PRIMENOM PEPELA IZ TERMOELEKTRANE KOLUBARA ................................................................................................................................................257

31. S.Abazi, I.Tomovski, P.Petrovski (Skoplje Makedonija) EKSPERIMENTALNA I NUMERIKA ANALIZA PONAANJA ARMIRANE ZEMLJE ....................................................................................................................................265

Tematska oblast 9 / Topic 9 DUBOKI ISKOPI I TUNELI DEEP EXCAVATIONS AND TUNNELS

32. N.Krstivojevi (Valjevo Srbija) ZATITA TEMELJNE JAME I SUSEDNIH OBJEKATA ZA IZGRADNJU STAMBENO POSLOVNE ZGRADE NA k.p. 1118/1 K.O. PALILULA U ULICI DALMATINSKA 14 U BEOGRADU ................................................................................................271

Tematska oblast 10 / Topic 10 STABILNOST KOSINA I KLIZITA SLOPE STABILITY AND LANDSLIDES

33. Z.Radi, V.Ili (Beograd Srbija) MERE ZA SPREAVANJE POJAVA NESTABILNOSTI KOSINA IZAZVANIH TEHNOGENIM PROCESIMA NA AUTOPUTU E-75 U SRBIJI ................................................................277

34. .ugi (Beograd Srbija) PRIMENA SOFTVERSKIH PAKETA FLAC I GEOSTUDIO ZA NASUTE BRANE I NASIPE ..........................................................................................................................285

35. M.Jovanovi, I.Vasi (Novi Sad Srbija) SANACIJA OBJEKTA NA KLIZITU DELIMINIM POTKOPAVANJEM I PODIZANJEM HIDRAULINIM PRESAMA U EROVIU ....................................................................291

36. A.Spahi (Srajevo BiH) RACIONALIZACIJA SANACIJE PLITKIH KLIZITA U ZAVISNOSTI OD IZBORA GEOMETRIJE ARMIRANOBETONSKE POTPORNE KONSTRUKCIJE .........................299

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37. A.Spahi (Sarajevo BiH) FUNKCIJA ZAVISNOSTI TROKOVA I FAKTORA SIGURNOSTI KOD SANACIJE PLITKIH KLIZITA U GLINAMA SA BETONSKOM POTPORNIM KONSTRUKCIJAMA .............................................................................................................307

38. B.Susinov, K.Lazarov (Skopje, Strumica Makedonija) STABILIZACIJA KOSIN OPTEREENE OBJEKTOM PRIMENOM ARMIRANOBETONSKOG ZIDA .................................................................................................................317

Tematska oblast 11 / Topic 11 HIDROTEHNIKI NASIPI I NASUTE BRANE FLOOD PROTECTION DYKES AND EARTH AND ROCKFILL DAMS

39. M.Vuini (Podgorica Crna Gora)

PRILOG SEIZMIKOJ ANALIZI NASUTIH BRANA ................................................................................325

Tematska oblast 12 / Topic 12 IPOVI, DIJAFRAGME I DRUGE TEHNOLOGIJE FUNDIRANJA PILES, DIAPHRAGM WALLS AND OTHER FOUNDATION METHODS

40. D.Mandi, A.Kikovi, M. Hranisavljevi (Beograd Srbija)

PRIMENA JET GROUTING TEHNOLOGIJE KOD FUNDIRANJA U DUBOKOJ TEMELJNOJ JAMI STAMBENO POSLOVNOG KOMPLEKSA NA NOVOM BEOGRADU ...................................................................................................333

41. P.Santra, .Baji (Subotica Srbija) ANALIZA VARIJANTE ZATITE TEMELJNE JAME I SUSEDNOG OBJEKTA ...................................343

42. A. Liolios, K.Liolios, B.Folic (Xanthi Greece, Belgrade Serbia) DYNAMIC PILE-SOIL INTERACTION UNDER ENVIRONMENTAL EFFECTS: NUMERICAL APPROACHES .......................................................................................................................349

43. .Rahimi (Mostar BiH) ISPITIVANJE BUENIH IPOVA PROBNIM OPTEREENJEM NA PRITISAK ....................................355

44. D.Zlatkov, M.Stanojev, S.Budi (Ni Srbija) VIEKRITERIJUMSKA OPTIMIZACIJA FUNDIRANJA OBJEKATA "MAGNETTO" KOMPLEKS "FIAT AUTOMOBILI SRBIJA" KRAGUJEVAC ........................................363

45. N.Davidovi, Z.Boni, V.Prolovi (Ni Serbia) GEOTECHNICAL CONDITIONS FOR THE FOUNDATION OF THE SCIENCE AND TECHNOLOGY PARK BUILDING IN NI ..................................................................371

Tematska oblast 13 / Topic 13 GEOTEHNIKA SAOBRAAJNICA: PUTEVI, ELEZNICE I AERODROMI GEOTECHNICAL ASPECTS OF ROADS, RAILWAYS AND AIRPORTS

46. M.Stevanovi, S.Bogdanovi (Beograd Srbija)

ANALIZA REZULTATA UPOREDNIH ISPITIVANJA MATERIJALA STABILIZOVANOG CEMENTOM I HIDRAULINIM VEZIVOM ..........................................................377

47. Z.Bai, A. Dananovi (Tuzla BiH) UPOTREBA NUS PROIZVODA PROCESA PROIZVODNJE SODE ZA IZGRADNJU DONJEG STROJA PUTEVA ..................................................................................................383

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Tematska oblast 14 / Topic 14 DEPONIJE VRSTOG OTPADA, EKOLOKI ASPEKTI GEOTEHNIKE ENVIRONMENTAL GEOTECHNICS, SOLID WASTE DISPOSAL

48. D.Raki, L.aki, S.ori (Beograd Srbija)

MEUZAVISNOST PARAMETARA STILJIVOSTI STAROG KOMUNALNOG OTPADA I KOEFICIJENTA POROZNOSTI ..................................................................391

49. S.okanovi (Beograd Srbija) GEOTEHNIKA ISTRAIVANJA ZA POTREBE PROIRENJA DEPONIJE KOMUNALNOG OTPADA U KRALJEVU .................................................................................................401

Tematska oblast 15 / Topic 15 MIKROZONIRANJE I SEIZMIKI RIZIK SEISMIC MICROZONING AND SEISMIC RISK

50. M.Vuini (Podgorica Crna Gora)

NEKI POJMOVI O ZEMLJOTRESIMA SA ASPEKTA POTREBA GRAEVINARSTVA ......................405 51. N.Mani, D.Luki (Novi Pazar, Subotica Srbija)

UTICAJ KVALITETA NASIPA NA SEIZMIKU POUZDANOST INFRASTRUKTURNIH OBJEKATA ............................................................................................................413

Tematska oblast 16 / Topic 16 OBRAZOVANJE U OBLASTI GEOTEHNIKE, SVI NIVOI OBRAZOVANJA KADROVA EDUCATION IN GEOTECHNICAL DOMAIN, ALL LEVELS

52. M.Hamova, G.Frangov, H.Zayakova, A.Mihailov, M.Periklijska

(Sofia Bulgaria) PROBLEMS AND FUTURE DEVELOPMENT OF GEOTECHNICAL TRAINING .................................421

53. J.Josifovski, M.Jovanovski, J.Papi, S.orevski, I.Peevski (Skoplje Makedonija) NEKE NOVOSTI U STUDIJSKOM PROGRAMU ZA GEOTEHNIKU ......................................................425

54. M.Trifkovi, .Nestorovi (Subotica, Kladovo Srbija) NEKI ASPEKTI STICANJA VETINA I ZNANJA U INENJERSKIM OBLASTIMA ..................................................................................................................................................431

Tematska oblast 17 / Topic 17 OSTALE TEME OD ZNAAJA, NEOUBUHVAENE TEMAMA OD 1 DO 16 OTHER TOPICS OF INTEREST NOT COVERED BY THE LIST ABOVE

55. A.Zahariev, G.Frangov, I.Zahivko (Sofija Bugarska)

INCREASE OF THE OVERALL STABILITY OF HPP ROSITZA 1, BULGARIA ............................................................................................................................435

56. S.Zdravkovi, D.Zlatkov, M.Stanojev (Ni Srbija) OCCURRENCE OF GREAT DIFFERENCE OF BENDING MOMENTS, DUE TO THE SETTLING OF THE SUPPORTS DEPENDING ON THE CROSS SECTION, CALCULATED ACCORDING TO THE SECOND ORDER THEORY ......................................................443

57. D.Raki, I.Basari, N.ui (Beograd Srbija) GEOTEHNIKI ASPEKTI ODRIVOG RAZVOJA ENERGETSKE GEOSTRUKTURE ......................455

58. M.Trifkovi, .Nestorovi, T.Milutinovi, G.Pejii (Subotica, Kladovo Srbija, Trebinje-R.Srpska-BiH, Brko - BiH) IZBOR TAAKA ZA DEFORMACIONU ANALIZU TLA I OBJEKATA PRIMENOM GEODETSKIH METODA ..............................................................................................................................463

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PREDGOVOR / (FOREWORD) Raznolikost geotehnikih uslova u naoj zemlji i socioekonomski odnosi u drutvu, kao i poloaj nae zemlje u svetu poslednjih godina, doveli su do zaostajanja za razvijenijim zemljama sveta u oblasti graevinske geotehnike. Zbog toga postoji potreba da se rezimiraju dosadanji rezultati i dostignua u ovoj vanoj oblasti u irokom spektru segmenata i to od metoda primenjenih geotehnikih terenskih istranih radova, laboratorijskih ispitivanja, primene savremenih teorijskih i numerikih postupaka, metodologije analize i projektovanja, kao i u oblasti praktine graevinske operative. Uspeno odrana savetovanja o geotehnikim aspekima graevinarstva (prvo Savetovanje na Kopaoniku 2005. godine, drugo Savetovanje u Sokobanji 2007. godine, tree i etvrto Savetovanje odrano na Zlatiboru 2009. i 2011. godine) podstaklo je Savez graevinskih inenjera Srbije (SGIS) da zajedno sa Srpskim drutvom za mehaniku tla i geotehniko inenjerstvo, uz podrku Ministarstva prosvete, nauke i tehnolokog razvoja Republike Srbije i Inenjerske komore Srbije, organizuje peto Savetovanje sa istom osnovnom tematikom. Osnovni cilj Savetovanja je razmena iskustva strunjaka razliitih profila i specijalnosti koji se bave geotehnikom. Savetovanje treba da ukae na glavne pravce razvoja ove struke koji bi odgovarali uslovima i potrebama u ovoj fazi izgradnje nae zemlje. Pored toga, to je prilika da se razmotri i stanje nae regulative u ovoj oblasti i potrebe njenog usaglaavanja sa najnovijim internacionalnim i evropskim standardima. Zbornik radova sa etvrtog Savetovanja sadri ukupno 58 radova koje je Programski odbor nakon pregleda prihvatio za izlaganje na Savetovanju. Na poetku zbornika tampani su radovi po pozivu istaknutih strunjaka, a ostali radovi su razvrstani u ukupno 17 tematskih grupa koje obuhvataju praktino sve aspekte geotehnike, i to.

1. NORMATIVI TEHNIKI PROPISI U GRAEVINSKOJ GEOTEHNICI U SVETLU USAGLAAVANJA SA EN STANDARDIMA

2. STANJE GEOTEHNIKE U NAOJ ZEMLJI 3. GEOTEHNIKA U PROJEKTOVANJU I IZVOENJU OBJEKATA INFRASTRUKTURE 4. GEOTEHNIKI ASPEKTI GRAENJA U URBANIM SREDINAMA 5. ISTRANI RADOVI, KARAKTERISTIKE TLA I STENA, KARAKTERIZACIJA I

KLASIFIKACIJA TERENA 6. MODELI GEOMATERIJALA I NUMERIKE METODE 7. PREDVIANJE I REZULTATI OSMATRANJA OBJEKATA, OPSERVACIONI METOD 8. POBOLJANJE TLA, ARMIRANJE, INJEKTIRANJE, DRENAE I DRUGO 9. DUBOKI ISKOPI I TUNELI 10. STABILNOST KOSINA I KLIZITA 11. HIDROTEHNIKI NASIPI I NASUTE BRANE 12. IPOVI, DIJAFRAGME I DRUGE TEHNOLOGIJE FUNDIRANJA 13. GEOTEHNIKA SAOBRAAJNICA: PUTEVI, ELEZNICE I AERODROMI 14. DEPONIJE VRSTOG OTPADA, EKOLOKI ASPEKTI GEOTEHNIKE 15. MIKROZONIRANJE I SEIZMIKI RIZIK 16. OBRAZOVANJE U OBLASTI GEOTEHNIKE, SVI NIVOI OBRAZOVANJA KADROVA 17. OSTALE TEME OD ZNAAJA, NEOBUHVAENE TEMAMA OD 1 DO 16

SGIS zahvaljuje ovim putem preduzeima i institucijama koje su pomogle odravanje ovog Savetovanja. SGIS takoe zahvaljuje lanovima Organizacionog odbora i Programskog odbora kao i autorima radova na uloenom trudu i njihovom stvaralakom radu u pripremi radova. Nadamo se i elimo da etvrto savetovanje SGIS o geotehnikim aspektima graevinarstva bude plodonosno i da se svi uesnici vrate u svoju sredinu obogaeni novim saznanjima i kolegijalnim poznanstvima.

EDITOR: Prof.dr Radomir Foli, Novi Sad

Beograd, oktobar 2013. TEHNIKI UREDNIK: Mr Aleksandar uki, Beograd

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UDK: 624.154.042.7 Pregledni (nauni) lanak

REVIEW PAPER

BOX-SHAPED DEEP FOUNDATIONS TO IMPROVE THE BEARING-SETTLEMENT BEHAVIOUR OF STRUCTURES

Em.O.Univ.-Prof. Dipl.-Ing. Dr.techn. Dr.h.c.mult. Heinz Brandl

Vienna University of Technology, Vienna, Austria

ABSTRACT Box shaped deep foundations consist primarily of pile walls or diaphragm walls (but also of deep/mixing walls or jet grouting walls). Combined systems of pile walls and jet grouting columns are also used. Such schemes have proved suitable for high-rise buildings, for bridges, silos, power stations, etc. Special applications are strengthening of old foundations (e.g. river bridges against scouring) and buildings in seismic zones. From theory, comprehensive model tests, and numerous site measurements and observations it could be concluded, that box-shaped deep foundations exhibit significant advantages over conventional pile or diaphragm panel foundations, because concrete elements and enclosed soil form a quasi-composite body with a high bearing capacity in vertical and horizontal direction. Moreover, they have a high resistance to earthquake, soil liquefaction and cyclic/fluctuating loading processes. The paper comprises theory, test results, design methods and case histories. KEYWORDS: Piled raft foundations, diaphragm wall foundations, box-shaped deep foundations,

earthquake resistance, scour resistance

SANDUASTI DUBOKI TEMELJI ZA POBOLJANJE PONAANJA NOSIVOST-SLEGANJE KONSTRUKCIJE

REZIME Sanduasti duboki temelji prvenstveno sastoje se od zidova formiranih pobijanjem ipova ili dijafragmi (ali takoe od dubokih/meovitih zidova formiranih mlaznim injektiranjem. Koriste se i kombinovani sistemi zidova od ipova i stubova dobijenih mlaznim injektiranjem. Ovaj nain je adekvatan za visoke zgrade, mostove, silose, energane, itd. Posebne primene su pojaavanje postojeih temelja, tj. mostova preko reka usled podlokavanja i zgrada u seizmikim zonama. Na osnovu teorije, opsenih eksperimenata na modelima i brojna merenja i monitoring moe se zakljuiti da sanduasti duboki temelji ispoljavaju znaajne prednosti u odnosu na klasine ipove ili temelj panel-dijafragmi jer betonski elementi uokviruju (zatvaraju) tlo kao kvazi-spregnuto telo velike nosivosti u vertikalnom i horizontalnom pravcu. Oni ak imaju veu seizmiku otpornost, pojavu likvefakcije i cikliko/ fluktuarijue procese optereenja. U lanku su opisani i analizirane teorijske osnove, rezultati eksperimenata, metode projektovanja i studije sluaja. KLJUNE REI: AB ploa na ipovima, temelji od zidova dijafragmi, sanduasti duboki temelji,

otpornost na zemljotrese, otpornost na podlokavanje/ispiranja

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BOX-SHAPED PILE AND DIAPHRAGM WALL FOUNDATIONS

Introduction

Box-shaped foundations have proved suitable for high-rise buildings, for bridges, silos, power stations etc. Special applications are foundations in creeping slopes (Fig. 1), strengthening of old foundations (e.g. river bridges against scouring, buildings in seismic zones (Fig. 2)). In principle all forms of ground plans are possible (Fig. 3). Box-shaped foundations act as a compound body consisting of piles (or diaphragm walls, deep-mixing walls or jet grouting walls) and the enclosed soil. This quasi-monolith can transfer high vertical and horizontal forces. Walls and capping raft form a box, which acts physically like a pot turned upside down. Consequently, the settlements are smaller than for conventional pile groups, and the earthquake resistance is significantly higher. Pile boxes (of bored or auger piles) represent a special form of piled raft foundations utilising the enclosed soil core as an integrated load transfer member. This is also the case, if diaphragm walls instead of pile walls are installed (Fig. 4). Intermittent pile walls with jet grouting columns between the piles are sometimes a cost-effective (and environmentally friendly) alternative to secant pile walls (Fig. 5). Thus, closed walls with a full shear bond can be obtained without excavating material for not reinforced (primary) piles; furthermore each pile can be reinforced. A certain disadvantage is the requirement of an additional site equipment. The optimum clear pile spacing lies typically between 0.2 to 0.5 m depending on the soil properties, required lengths of piles and jet grouting columns, on the jet grouting technique and on static requirements.

Figure 1. Box-shaped pile foundation for a bridge pier in unstable slope. Uphill pile wall tied back with prestressed anchors as additional safety measure.

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Figure 2. High earthquake resistance of box-shaped deep foundations. The confinement of the ground enclose by pile walls or diaphragm walls reduces the soil deformation below buildings significantly.

Figure 3. Box-shaped pile foundation for a river bridge pier (scheme proven also for silos).

Figure 4. Box-shaped foundations on slurry-trench walls (diaphragm walls) for a river bridge pier.

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Figure 5. Scheme of combined wall system for deep box foundations: Reinforced bored (or auger) piles and jet grouting columns in between.

Model tests Comprehensive model tests were performed to investigate parameters influencing the bearing-settlement behaviour of box-shaped pile foundations. The research program comprised 70 tests including the following test series (Hofmann 2001, Brandl 2001, Brandl & Hofmann 2002): Pile boxes with inner piles (according to general design practice); Pile boxes without inner piles; Pile boxes without soil infill (simulating zero-stiffness of the enclosed soil); Pile boxes filled with concrete (simulating a monolithic block); Conventional pile groups (axial spacing a 2d); Close contact or free gap between raft and soil beneath; Single piles.

Figure 6. Two examples of box-foundations used for the standard model tests (scale 1:50). Equivalent diameter D for a circular box-foundation for box S.

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Figure 7. Dimensionless load-settlement curves for the pile box S. Pile length l = 40 cm. s = settlement, Q = total load on the foundation, d = pile diameter, A = foundation area (cross sectional area within circumference of pile

box), = density of soil.

Figure 8. Similar to Fig. 37, but pile box L. Tests with conventional pile groups and single piles were conducted to compare the load-settlement behaviour of the different pile patterns. Furthermore, pile diameter (d), pile length (l) and density of soil () were varied to check their influence. The standard test series were performed with uniform quartzitic sand: d50 = 0.75 mm, dmax = 2 mm, Cu = 3. Figure 6 shows two standard types of investigated box-foundations. The pile pattern was similar to the design of foundation alternatives for a long river bridge. During the tests the load-settlement curves until failure, the settlement troughs and the pile forces in five or six measuring levels were registered. Figures 7 and 8 show some test results in normalized diagrams. The data are given dimensionless to enable a direct comparison with results gained for conventional pile groups or from in situ measurements on construction sites. Moreover, dimensionless diagrams can be applied more easily to larger scales. The diagrams demonstrate the effect of pile arrangement and intensity of bond within the pile box. The installation of inner piles reduces the settlement, which can be expressed by a small boxes is relatively greater, whereby, of course, large box-foundations as a whole can take higher total loads due to their larger area and pile number. Figure 9 illustrates the composite effect: The hatched zone between pile box without infill and full monolith depends on the bond factor.

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The influence of the ratio of box area A to box circumference U increases with the limit pile load. This ratio corresponds to the hydraulic radius R = A/U. Hence, boxes with a small hydraulic radius (i.e. long-stretched) can transfer higher loads than those with a square or circular shape. This coincides well with the theory, because square or circular foundations cause a higher stress concentration in the ground.

Figure 9. Influence of stiffness of enclosed soil or bond effect between piles and enclosed soil of the box-foundation S; pile length l = 40 cm.

Figure 10. Box-factor of deep box-foundations versus ratio A/U/d, where A = area of pile box, U = circumference of pile box, d = pile diameter. Derived from model tests.

The portion of external load directly taken by the soil core increases with increasing hydraulic radius of the pile box, assuming a similar pile arrangement. From the model tests, a box-factor could be deduced:

totalsoil QQ = , (1) hence totalpile QQ )1( = , (2) whereby 4.00 , (3) The box-factor increases with the stiffness of the soil core. Usually it lies below 0.4. A higher value can be obtained if the soil core is improved by jet grouting or deep mixing. Figure 10 shows the box-factor for limit loads, which characterise a beginning steepening of the load-settlement curve. The -lines should not be extrapolated linearly to values higher than A/U/d = 2. It is rather recommended to

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design piles with a box-factor that then is kept constant (for safety reasons). When approaching failure load the forces concentrate in the piles, because the ratio of stiffness of piles to plastified soil increases. But due to a self-regulating behaviour box-foundations do not have a clear ultimate load. Figure 10 demonstrates the great influence of the cell size(s) on the load transfer via the soil core(s). The portion of external load directly taken by the enclosed soil of the box-foundation increases with the hydraulic radius A/U or A/U/d. A cohesion of the soil has no significant effect on the ratio Qsoil/Qtotal, but it influences the load transfer mechanism of the piles, hence the percentage of skin friction force and base resistance force. Figure 10 represents only one among various correlations because the box-factor depends on a series of parameters: Ratio A/U/d; Slenderness of the box-foundation, l/D; Ratio of stiffness of concrete members (Econcrete) to soil (Esoil); Multi-cellular pattern of the box foundation; Ratio of service load to limit or rupture load; Settlement. The portion of external load directly transferred from the raft into the soil (Qsoil/Qtotal) decreases with pile length l and box slenderness l/D respectively. The main reduction occurs between l/D = 0 (i.e. flat foundation where the raft takes 100 % of Qtotal) and l/D = 0.5 to 0.75 where the raft usually takes about 60 to 30% of Qtotal.

Figure 11. Base pressure of the piles versus settlement of the pile box S (Fig. 36). Pile length l = 50cm. The transfer of vertical loads by a box-shaped pile foundation concentrates rather on the inner piles than on the outer ones. This effect increases with increasing total load (e.g. Fig. 11) and is caused by a silo pressure within the cells (Brandl 2001). The base pressure of the piles increases with the size of the soil cells because larger cells facilitate higher silo pressures. In the upper zone of the box, the ring walls are subjected to a lateral earth pressure difference that is directed outward. With increasing depth, the horizontal silo pressure is widely compensated by the earth pressure at rest acting on the outer face of the box-foundation. Therefore, adjacent piles should exhibit sufficient bond along their connecting line (mainly in the upper zone), i.e. secant piles are advantageous over tangent piles. In the case of contiguous or even intermittent pile walls, a load transferring closure can be obtained by jet grouting between the spandrels. The effectiveness of piles forming cross walls in a deep box-foundation can be quantified by dividing the settlement reduction by the increase of the proportional pile number when adding inner piles to form a multi-cellular pattern. The model tests exhibited that piles forming cross walls in long-stretched boxes have a larger effect than those in square-shaped or circular cells.

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Theory and calculation General

In conventional design practice, the bearing capacity of the capping raft of deep foundations is usually neglected. But box-shaped deep foundations behave as a compound body: the enclosed soil cannot move laterally and takes part in bearing external loads. Consequently, the capping raft can be designed to take a signifyc cant percentage of the forces from the structure above by transferring it directly into the ground. Comprehensive model tests and in-situ measurements have shown that the settlement of such box-foundations is smaller than it would be in the case of conventional groups of piles or diaphragm wall panels. Avoiding the lateral deformation of the soil core and minimizing its shear deformation leads to a significant reduction of settlements, because the foundation system acts like a pot turned upside down. For the design and calculation of such deep box-foundations, several hypotheses have proved suitable: Half-space hypothesis Limit case hypotheses Subgrade reaction models Numerical models.

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Figure 12. Design charts for deep box-foundations. Settlement curves for a cylindrical box foundation under a unit load of Q = 1 kN. Slenderness l/D of the foundation as parameter, whereby l = depth of pile or diaphragm wall

foundation. Unit modulus of soil Es = 20 MN/m, d = wall thickness. For non-cylindrical foundations: D = equivalent diameter.

Elastic-isotropic half-space hypothesis Figure 12 shows a design chart for the determination of the unit-settlement of a cylindrical box-foundation depending on its diameter and slenderness. It is based on the half-space hypothesis (Brandl 1987, 2001). Originally, the integration of Mindlins equations was per performed for a circular diameter referring to the axis of the

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circumference wall (Hazivar, 1979). Therefore, if the box has a rectangular or polygonal shape, an equivalent diameter must be chosen (see also Figure 29). The theoretical diameter should be somewhat smaller than the outline of the cell (e.g. minus d/2), depending on the pile spacing (intermittent, contiguous or secant). This is an allowable approximation that has proven suitable in practice, especially for commonly designed and utilised rectangular box-foundations. In the case of a rectangular foundation box, the transformation into an equivalent diameter means a theoretically stronger stress concentration especially in the case of long-stretched boxes (e.g. Figure 6, Box S). This effect justifies an equivalent diameter somewhat larger than the axial wall spacing and fits better to that area where friction forces are transferred in reality. Single elements within the enclosed soil core reduce the settlement, but not significantly. Transverse walls have a greater effect. Another purpose of such additional inner elements is to stiffen the foundation-box and to gain a statically optimal support for the capping raft. Furthermore, the bearing capacity for horizontal loads and moments increases, and the earthquake resistance is improved significantly.

Figure 13. Cell-factor c of (multi-cellular) deep box-foundations versus the ratio A/U/d. Number of cells, n, of the

box-foundation as parameter. The settlement assessment curves of Figures 12 and 27 were developed for cylindrical box-foundations without stiffening elements inside. But comprehensive model tests on box-shaped pile foundations with and without inner piles disclosed that the installation of inner walls increases the bearing capacity and reduces the settlement. From model tests and in situ measurements on numerous sites a cell-factor c could be deduced, which describes the effect of a multi-cellular shape of the box-foundation (Figure 13). It was determined for service loads corresponding to about 50% of the limit loads. Commonly it varies between

0.15.0 c , (4) The maximum value occurs if no inner piles are installed, the minimum refers to a multi-cellular pattern with relatively small cells. In the latter case the pile (or diaphragm wall) foundation behaves increasingly like an idealized quasi-monolithic block foundation with a deep-lying foundation base.

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Figure 13 illustrates that the cell-factor depends on the hydraulic radius A/U of the box foundation, on the pile diameter d (or wall thickness d), and on the number of cells within the box. The relatively greatest settlement-reducing and stiffening effect is gained with two or three cells. Usually, large foundations should exhibit at least three cells. The theoretical minimum of c is obtained if the entire box is filled with concrete elements (or jet grouting columns or deep mixing columns). But this is uneconomical. A cost-effective compromise, however, could be a local soil core improvement by (jet) grouting. Nevertheless, experience has shown that the cell-factor used for practical settlement assessment should not be assumed smaller than c = 0.5. From Figures 12 and 13 the settlement s of a box-shaped deep foundation with an equivalent diameter D can be calculated as follows:

zEE

QQs

s

stotc

= 1,

1

, (5)

c = cell factor of the deep box-foundation (from Figure 13)

totQ [kN] = settlement-effective total load on top of the pile group (or diaphragm wall group)

1Q = unit load, i.e. 1Q = 1 kN

sE [kN/m] = modulus of soil (mean value)

1,sE = unit modulus, i.e. = 20 MN/m

z = unit settlement from Figure 12 Equation (5) is primarily valid for a wall thickness of about d = 1m, but may be used for d = 0.8 to 1.5m with sufficient accuracy. (In the case of diaphragm walls even for d = 0.6 m). It is strictly speaking based on a Poissons ratio of = 0.3 and on a modulus ratio of structural members to soil of about 103. But values of 0.2 < 0.5 have no relevant influence on the result. Furthermore, a variation of the ratio Epile : Esoil between 5.102 to 5.103 is allowable if the soil modulus is properly chosen. Hence equation (5) and Figure 12 have proved suitable for a wide range of different soils. Only for very soft clays too large settlements are calculated, and for very stiff overconsolidated clays the cell-factor should be neglected (hence approximately c = 1 also for multi-cellular boxes). Furthermore, this formula is usually limited to soils with a modulus of about Es 100 MN/m2.

Limit case hypotheses Limit case hypotheses and analyses refer to theoretically idealized limit assumptions (upper and lower bound approaches) and are not necessarily identical with limit load analyses or ultimate load conditions of deep foundations. For assessing the bearing capacity of box-shaped foundations, two methods have proved successful in design practice: Calculating the bearing capacity of the single pile or single diaphragm element (Figs. 14, 15) safety factor

F1. Calculating the bearing capacity and settlement of the box-foundation as a quasi-block according to the

monolith theory (Fig. 16) safety factor F2.

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Evaluating the bearing capacity of single elements provides only fictitious limit case values because the bond effect between concrete elements and enclosed soil core is neglected. Thus, maximum pile or diaphragm wall loads are calculated. But actually, single elements cannot fail because of the composite effect and the rigid (reinforced) connection of the piles or diaphragm panels with the capping raft. Moreover, deep box-foundations exhibit a self-regulating bearing behaviour, especially if the boxes are stiffened with inner walls: in the case of a local overloading of the soil around a pile, stress redistribution is possible.

Figure 14. Scheme of load transfer in a box-shaped deep foundation with inner pile- or diaphragm walls.

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Figure 15. Subgrade reaction model for a strengthened foundation of the central river pier of an old Vienna Danube bridge showing soil responses to superimposed loads. Box-shaped new pile foundation consisting of secant piles and soil improvement in the upper part. Hatching on the diagram (below) indicates the difference between actual

settlement and idealised model in the case of load increase.

Figure 16. Box-shaped foundation (consisting of bored piles or diaphragm walls and the enclosed soil core) loaded by vertical and horizontal forces and moments: Idealised model quasi monolith of the limit case hypothesis for

determining the safety factor F2 against ground failure and evaluating the settlements.

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Therefore, very low safety factors are sufficient for this theoretical model: usually F1 1.15. For short construction stages or catastrophic conditions even values of F1 = 1.05 have been allowed. Contrary to the monolith-theory, skin friction may be taken into consideration along the outside and inside face of the foundation-box, but not between the single elements. The other limit case hypothesis is an idealised monolith-theory. According to Fig. 16, a full bond effect between deep foundation elements and the closed soil is assumed. This compound body comprises the outer circumference of the foundation if secant piles or diaphragm walls are installed. In the case of contiguous piles, the theoretical area should be reduced by at least half a pile diameter. For the quasi-monolith, only skin friction along the outside surface of the foundation box may be taken into account. The monolith-theory provides minimum pile or diaphragm wall loads. However, a full composite effect occurs only theoretically but hardly in practice. Therefore, relatively high safety factors are required: about F2 3.0 if conventional calculation methods for evaluating the base failure of equivalent shallow foundations are used. Short-term traffic loads do not reach the toe of the deep box-foundations but are more or less directly transferred into the upper soil zone, unless the box has an exclusively end-bearing character. For settlement analyses, the monolith-theory has proved practicable and sufficiently accurate in engineering practice by assuming the base of the box-foundation as the fictitious surface of the half space. The theoretical contact pressure includes the reduction of the total load Q by the skin friction Qs.

Case histories General

Numerous data from in-situ measurements have been collected over a period of about 35 years. They comprise stress and deformation/settlement measurements of box-shaped deep foundations of bridges, hydropower plants, industrial buildings and high-rise buildings. The ground plan of the box-foundations was rectangular, circular, elliptical or polygonal and mostly stiffened by transversal and/or longitudinal wall elements. Sometimes single piles or diaphragm wall panels were additionally installed within the cells (for static reasons; to compensate installation failures, etc.). The ground conditions varied from very soft clay to stiff overconsolidated clay, from loose to dense sands, gravel, heterogeneous colluvium, from weathered slope deposits to decomposed rock.

Figure 17. Ground plan of a box-shaped foundation for a highway bridge in silty river sediments. Bored piles,

diameter d = 0.9 m. Single piles (hatched) only additional to overcome construction difficulties and local inhomogenities in the subsoil. Black piles: reinforced; white piles: not reinforced.

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The wall systems and the way of installation were also different. Both have an influence on the load-settlement behaviour of the box-shaped foundations. Diaphragm walls, for instance, provide a better transfer of shear forces between the concrete panels than contiguous pile walls, but on the other hand have frequently a smaller skin friction. The in-situ measurements confirmed that the percentage of load taken either by the capping raft or by the piles (or diaphragm walls) depends on various parameters, such as: Cross section (incl. pile pattern etc.) and slenderness of the foundation-box; Ratio of stiffness of concrete elements and soil; Magnitude and distribution of external loads (V,H,M); Ratio of service load to ultimate (failure) load; Ground properties; Vertical and horizontal soil displacement; Magnitude and distribution of the contact stress between raft and soil; Foundation depth; Depth of excavation (construction pit); Installation factors.

Highway bridge, Austria Consequently, the results of in-situ measurements scatter relatively widely, including several changes also during the construction period. In the following, a case history is selected which represents rather weak soil conditions. The piers of a highway bridge had to be founded in deep-reaching young river sediments: Sandy gravel of about 4 m thickness, underlain by weak silts (sandy to clayey); the natural water content varied between the plastic and liquid limit, the dry density was d = 1,6 - 1,7 t/m. Figure 17 shows the pile arrangement, thus forming a box-shaped foundation. The enveloping piles are secant, therefore only every second one is reinforced. The interior piles are throughout reinforced and improve the load transfer from the bridge pier to the deep foundation. The construction was exe cuted in the year 1971. Nowadays the inner piles would be rather installed in a secant form to increase the stiffness of the box.

Figure 18. Partial view of a power station founded on a box-shaped arrangement of diaphragm walls (slurry trench

walls) in soft clayey silt.

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For the bridge design, the (differential) settlements were assessed rather cautiously, because this was one of the first box foundations. The measured value of s = 55 mm for the central bridge pier lies clearly below the prognosticated maximum of s = 100 mm, but coincides very well with the result derived from Figure 12 and Equation (5): ground area of the box-foundation, A = 155.6m circumference of box-foundation, U = 50.7m equivalent diameter of a cylindrical foundation, D = 14m pile length, l = 11m pile diameter, d = 0.88m slenderness of the foundation box, 1 : D = 0.8 total load (life load reduced), Qtot = 55000 kN modulus of subsoil (mean value), Es = 8 MN/m From Figure 13 a cell factor of c = 0.5 is derived (for 6 cells) and provides a unit settlement of z = 0.810-3 mm. This leads to a total settlement of

= 3108.0

820

1550005.0s mm55s = , (6)

Hydro-power plants

Figure 18 shows the ground plan of a hydro-power plant in the South of Austria. It comprises an operation hall, two pier-power houses and three weirs. The equipment consists of Kaplan turbines with vertical shafts, the machinery being extremely sensitive towards differential settlements. The entire power station is situated in a flood plain. The subsoil consists of river deposits (sandy gravel of medium density), which are underlain by soft fine-graded sediments. Though silty clays pre-dominate, sandy silts or uniform sands are locally embedded too. Accordingly, the soil characteristics scatter widely: natural water content wn = 15 - 50% dry density d = 1,3 - 1,7 t/m liquit limit wl = 20 - 60% plasticity index Ip = 0 - 30%

Figure 19. Cross section (in river flow direction) through a partition pier of the river Danube hydropower plant in

Vienna (14 000 m3/s flood capacity). Box-shaped foundation on diaphragm walls arranged as stiffening cells.

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The stiffness of the weak clays and silt increases with depth: Moduli of Es = 4 - 10 MN/m near the surface of the sediments and Es = 10 20 MN/m in a depth of 30 40 m. Standard penetration tests showed values of N30 = 1 to 10 in 4 to 36 m depth below the original ground. Though the base of the power station lies clearly beneath the original surface (i.e. in the soft sediments), and the additional load on the ground is relatively small, a deep foundation was unavoidable. It was designed as a box foundation of longitudinal and transversal diaphragm walls (0.8m thick) in order to: reduce differential settlements; increase the safety factor against earthquake (the power station is situated in an active seismic zone); prevent an unallowable under seepage of the power station. No open joints were designed between the power houses and the weirs. But the connections were reinforced in such a way that in the case of differential settlements they serve as a hinge.

- Figure 20. Vertical section through a box foundation with raked piles for a pier of a river bridge.

The deformation behaviour of the power plant coincided very well with the design prognosis: Due to the deep excavation, the subsoil heaved during the first construction stages. With increasing load, settlements occurred also depending on the required ground-water lowering in the construction pit. The total settlement finally has reached a maximum of s = 55 mm. Because of the rigid box-shaped foundation and structure, the differential settlements are practically negligible. Though the large scale foundation of this power station cannot be compared with a cellular box-foundation of bridge piers or high rise buildings, the diagram of Figure 12 provided even for this case an appropriate settlement assessment: Cross sectional area (ground area) of the of the effective box-foundation, A = 3000m equivalent diameter of a cylindrical foundation, D = 61m

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depth of diaphragm walls (mean value), l = 16m slenderness of the foundation box, 1 : D = 0,3 settlement-effective load (reduced by excavation masses), Qtot = 270000kN modulus of the subsoil (mean value beneath the diaphragm wall bases), Es = 15 MN/m From Figure 12 or 27 a unit settlement of z = 0,3.10-3 mm is derived and from Figure 13 cell factor of c = 0,5. This leads to a total settlement of

= 3103.0

1520

12700005.0s mm54s = , (7)

Figure 19 shows a cross section through the hydropower plant Vienna at the river Danube (designed for 14 000 m3/s, width = 450 m). This is the first facility constructed in a densely populated area, situated on a geological fault and in a seismic area. Consequently, a deep box-shaped foundation consisting of slurry trench walls forming stiffening cells was executed. Such foundation schemes have proved suitable for all kinds of buildings with high vertical and/or horizontal loads.

Figure 21. Horizontal section through the pile heads of Figure 50. Position of pile toes indicated by broken line

circles.

Figure 22. Horizontal sections through a box foundation with raked piles: Example of a small box. Section 1-1 is in

the level of pile heads, section 3-3 in the level of pile toes.

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Box-shaped foundations with raked piles

In rivers with shipping, high flow velocity and danger of scouring bridge piers (including their foundation) should be as small as possible and hydraulically friendly. Consequently, truncated cone box foundations are an interesting alternative to the classical prismatic shape. Figure 20 shows a vertical cross section of such a foundation, and Figure 21 illustrates the horizontal sections on top and toe of the piles. The Figures 22 and 23 show a smaller box foundation. In both cases the piles were excavated through a fly ash body which was filled into the box of precast elements placed on the river bed (and fixed by the pilot piles). This measure facilitated a precise installation of the raked piles and increased the composite behaviour of the box foundation. The pilot piles were installed before (from ships) in order to have fixing points for further equipment and for sinking the r.c. precast elements on the river bed. Therefore the steel casing was not withdrawn near their head zone. Moreover, the pilot piles were excavated to a greater depth than the standard piles in order to gain detailed additional information about the in-situ ground properties beneath the bottom of the box foundation. The installation of raked piles for a box foundation like a frustum of a cone or pyramid is rather difficult, especially if they have to be excavated from a ship or swimming platform. Consequently, but also for geometrical reasons, the box effect decreases with depth. Such a foundation may be considered a structure behaving between a box-foundation and a piled raft foundation.

Figure 23. To Figure 22: View of the longer side of the pile box.

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Beska Bridge, Serbia The Beska Bridge crosses the river Danube carrying the E-75 highway between Novi Sad and Belgrade. The new bridge had to be erected parallel to the existing one (from the early 1970s) and was opened in 2011. The total length of the bridge is 2005 m; the central part of the new structure consists of 5 spans (60+105+210+105+60 m) with a prestressed concrete hollow box girder of changing height. This Main Bridge was founded on a box-shaped pile system (Fig. 24), whereas the approach bridges rest on conventional pile groups. The main reasons for the box-foundation were as follows: Minimisation of (differential) settlements of the nearby existing bridge during the construction phases and in

the long-term. Minimisation of (differential) settlements of the new structure. High resistance against possible effects from riverbed erosion (deep scouring). High resistance against seismic impacts. High resistance against ship impact and excessive ice pressure.

Figure. 24. Beska Bridge, Serbia. Foundation of old structure on caissons, of new structure on box-shaped pile arrangements. Also indicated is the riverbed shaping against deep scouring at river pier No. 43.

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Figure. 25. Cross section to Fig. 24. To cope with these requirements the new foundation had to be deeper than the existing caisson of the old bridge. Additionally, a possible riverbed erosion was considered by neglecting the topmost 6 m of pile-box embedment in the geotechnical/static design. From Fig. 25 it can be seen that the minimum distance between old and new foundation was only 2.5 m. Therefore, an interaction between both bridges was unavoidable and required an extremely cautious installation of sheet piles and large diameter bored piles (d = 1.2 m). Within the quaternary sediments and the top zone of the tertiary sediments the spaces between neighbouring piles were grouted to provide a full shear force transfer along closed walls, and to increase skin friction (Fig. 25). Moreover, the concreting phases /speeds for the piers and superstructure had to consider the results of continual monitoring of the old and new structure. Fig. 26 shows two construction phases. At the beginning of the pile work pilot piles were excavated 3 m below the designed base of the box-shaped deep foundation to additionally investigate the subsoil at a large scale. Moreover, the excavated soil was continually checked during the entire pile installation (visual inspection, SPTs, soil sampling and laboratory tests). Thus, the in situ ground properties could be compared in detail with the design assumptions. If necessary, local adaptions were performed. All those measures led to differential settlements, which remained clearly below the allowable limit value of 10 mm, and confirmed again the advantages of box-shaped pile foundations under such challenging conditions: For the main piers of the new bridge settlements of s = 15 to 25 mm were predicted; the measured values were s 20 mm. For the main piers of the old bridge settlements caused by the new bridge were predicted between s = 7 to 12 mm. Measurements showed s 8 mm, whereby 5 mm occurred during pile installation. Consequently, it was not necessary to activate the contingency plan comprising the additional installation of Gewi-piles as "settlement brake".

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Figure. 26. Construction views to Figs. 24, 25. Left: Placing the

concrete for the scalloped guide wall for the piles and the temporary working platform. Right: Concreting the shaft of river pier No. 43.

COMPARISON OF DEEP FOUNDATIONS

In order to prove the reliability of geotechnical theories and the general application of test results to the practice, in-situ measurements on construction sites and on completed structures are essential. Figures 27 to 28 show some examples of deep foundations in tertiary sediments, overlain by quaternary river deposits (in Vienna and Lower Austria). The tertiary layers are over-consolidated and consist of sandy to clayey silt (locally silty sand and silty clay). The ground properties, of course, scatter in spite of the same geological genesis along the river Danube in Vienna and nearby. Nevertheless, the site conditions of the structures can be fairly compared. The deep foundations elements were large diameter bored or auger piles (d=0.9 to 1.2 m) or diaphragms walls (thickness, d=0.6 to 0.8 m). The data are again plotted in dimensionless diagrams similar to the results from the model tests. The normalized graphs enable a direct comparison.

Figure 27. Normalized load-settlement behaviour of some high-rise buildings in Vienna, founded on diaphragm

walls. s = settlement, d = thickness of diaphragm wall, Q = total load on the foundation, A = foundation area (horizontal sectional area within circumference of diaphragm wall box), = density of soil, H = height of building,

l = depth of diaphragm wall

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Figure 27 shows normalized load-settlement curves of several box-shaped foundations with diaphragm walls. A comparison underlines the following interacting influence factors: Length (depth) of wall elements: Widely superimposed by other factors, because l = 18 22 m is roughly of a

similar magnitude. Hence of secondary influence in this special case. Level of foundation head: The high-rise building without deep basement (Mischek Tower) settled more than

the others. Way of installation: The installation from a higher working level improves the bearing- deformation behaviour

of diaphragm walls (e.g. IZD-Tower) - and bored or auger piles likewise. The positive effect of an uncast pile length or uncast diaphragm wall depth could be observed on many construction sites. It is achieved because the soil along the top zone of the deep foundation is less disturbed during the installation procedure.

Influence of geological overconsolidation: The Twin Tower is situated in an area of less overconsolidation than the other buildings.

Arrangement of diaphragm walls: The buildings of the UNO-City Vienna are founded on box-shaped as well as on cross-shaped diaphragm wall elements. The latter settled more, because the composite effect between single barrettes or cross-shaped concrete panels and soil is clearly smaller than for box-foundations.

Figure 28, finally, compares high-rise buildings and bridges, whereby the following influence factors played a relevant role: Load area and total load, hence depth and extent of stress bulb in the ground: Millennium Tower and UNO-

City create a deep reaching stress field. Magnitude of geological overconsolidation: This is greater for both bridges (crossing the river Danube in

Lower Austria) than for the high-rise buildings in Vienna. Stiffening inner pile walls within a box-shaped foundation: The Tulln Bridge has piers with both options. Length (depth) of piles or diaphragm walls. Influence of local soil improvement: The top zone of the soil was improved by vibroflotation (Millennium

Tower) or by jet grouting (Pchlarn Bridge, where grouting was conducted within the piled box foundation). Beneath the Millennium Tower the surface of the tertiary sediments lies deeper than on the other sites. Therefore, vibroflotation of the top zone was very effective here and made shorter piles possible.

Composite effect of box-shaped foundations: The Millennium Tower has a piled raft foundation, the Tulln Bridge box-shaped pier foundations with and without stiffening inner pile walls.

The comparison of the load-settlement behaviour of several buildings illustrates very clearly the interaction of more or less relevant influence factors. Therefore, a reliable settlement prognosis can be achieved only by taking into account all these aspects. This means, that theoretical calculations have to be combined with empirical factors gained by in situ measurements (and model tests). Box-shaped deep foundations provide not only smaller settlements than comparable conventional pile or diaphragm wall foundations. They also make stress rearrangement easier, thus reducing stress constraints in the capping raft or beam and in the rising structure.

23

Figure 28. Normalized load-settlement behaviour of bridges and high-rise buildings, depending on several factors. Bridges and Millenium Tower rest on piles. Notations see Fig. 27, whereby also d = pile diameter, l = pile length.

CONCLUSIONS

Since the early 1970s box-shaped deep foundations have proved most suitable, as long-term measurements and site observations have confirmed. The scheme comprises a box-shaped arrangement of (bored or auger) piles or diaphragm wall panels, deep-mixing columns or jet grouting columns. From theory, comprehensive model tests, and numerous in situ measurements and observations it can be concluded that box-shaped deep foundations exhibit the following advantages: Concrete elements and enclosed soil form a quasi-composite body with a high bearing capacity. Transfer of high, concentrated loads at smaller total and especially differential set settlements than in the case

of conventional groups of piles or diaphragm wall panels. Smaller foundation area required than in the case of conventional pile groups with axial pile spacing of a 2d

to 3d (d = pile diameter). This is especially important for bridge piers situated in rivers. High resisting moment against lateral forces from high embankments on soft soil (acting on bridge abutments)

or from unstable slopes (e.g. creeping pressure). Very suitable in the case of strongly heterogeneous and anisotropy ground. High resistance to earthquake and soil liquefaction. High resistance to scouring (foundations in rivers, torrents, harbours). Very suitable in the case of dynamic loading processes: E.g. ship impact, dynamic loads due to waves and

unstable currents, (turbulent) wind loading of tall structures, shock loads due to unstable silo flow. Suitable for post-strengthening existing buildings, for instance piers of river bridges: Increase of stability

against ground failure in the case of scouring. Very suitable in the case of fluctuating and cyclic loading processes: E.g. cycles of (ground-)water level,

storage level fluctuations (oil tank farms, storage silos) or wave induced cyclic loads.

24

Figure 29. Settlement curves for a cylindrical box foundation under a unit load of Q = 1 kN. Slenderness l/D of the

foundation as parameter, whereby l = depth of pile or diaphragm wall foundation. Unit modulus of soil Es = 20 MN/m, d = wall thickness. For non-cylindrical foundations: D = equivalent diameter. See also Fig. 12.

REFERENCES Brandl, H. 1987. Deep box-foundations with piles and diaphragm walls in weak soils. Proceedings of 9th Southeast

Asian Geotechnical Conference. Bangkok. Brandl, H. 2001. Box-shaped pile and diaphragm wall foundations for high loads. Proceedings of the 15th

ICSMGE. Istanbul. Brandl, H. & Hofmann, R. 2002. Tragfhigkeits- und Setzungsverhalten von Kastenfundierungen. Research Report,

Volume 528. Federal Ministry for Traffic, Innovation and Technology. Vienna. Hazivar, W. 1979. Tragverhalten von Brunnengrndungen, Volume 14, Institute for Soil Mechanics and

Geotechnical Engineering, Vienna University of Technology. Hofmann, R. 2001. Trag- und Setzungsverhalten von Pfahlksten. Doctoral Thesis, Vienna University of

Technology. Japanese Geotechnical Society, 1998. Remedial measures against soil liquefaction. A.A. Balkema, Rotterdam. Mindlin, R.D. 1936. Force at a Point in the Interior of a Semi Infinite Solid. Physics, Vol. 7.

25

UDK: 624.131.3 Izvorni (nauni) lanak

ORIGINAL SCIENTIFIC PAPER

A STATIC LOADING PILE TEST AT THE SERVICEABILITY LIMIT STATE Ludvik Trauner, Andrej trukelj, Mirko Punder, Borut Macuh University of Maribor, Faculty of Civil Engineering, Smetanova ulica 17, 2000 Maribor, Slovenia; E-mail: trauner@uni-mb.si, andrej.strukelj@uni-mb.si, mirko.psunder@uni-mb.si, borut.macuh@uni-mb.si

ABSTRACT A static vertical and horizontal loading test of a pile was performed when constructing the covered excavation 8-1 on the under section of the highway PonikveHrastje, Slovenia. The engineering structure is located on the section at about 1600 km of the deviation 1-12 of the regional road. Besides the standard testing procedures with vertical loading, the horizontal loading test was performed and additionally there was new monitoring technology based on specially developed strain sensors installed inside the pile body. Along the four verticals circularly beside the longitudinal reinforcement bars of the pile, 32 specially developed sensors were introduced. On the basis of the measured results the normal strains along four verticals at the distance of three-quarters of the radii of the pile from the pile axis were measured. Taking into consideration the elastic modulus of the concrete the normal stresses in the axial direction of the pile were also calculated and afterwards the shear stresses along the pile shaft were estimated as well as the normal stresses below the pile toe. The estimation was made by considering a constant value for the pile diameter. The measured results were also compared to a computer simulation of the pile and the soil behaviour during all the successive test phases. The strain measurements inside the pile body during the vertical and horizontal loading tests in the present case did not have the purpose to develop an alternative method of pile loading tests. It gave, in the first place, the possibility of a closer look at the strains and stresses of the most unapproachable parts of different types of concrete structure elements, especially the piles and the other types of deep foundations. The presented monitoring technology proved itself to be very accurate and consistent. KEY WORDS: piles, deep foundations, static loading test, strain measurement technologies, elasto-

plastic modelling, finite-elements method

TEST STAIKOG OPTEREENJA PRI GRANINOM STANJU UPOTREBLJIVOSTI

REZIME Statiki test za vertikalno i horizontalno optereenje ipa je obavljen kada je graen pokriva iskopa 8-1 ispod preseka autoputa Ponikve Hrastje u Sloveniji. Inenjerski objekat je lociran u preseku oko 1600

26

km devijacije I 12 regionalnog puta. Pored standardnog postupka ispitivanja pod vertikalnim optereenjem izveden je i test pod horizontalnim optereenjem za koji je koriena nova tehnologija osmatranja zasnovana na specijalno razvijenim senzorima registracije dilatacija instalisanih izvan tela ipa. Podu etiri vertikale kruno pored podunih ipki armature ipa uvedena su 32 razvijena senzora. Na bazi rezultata merenih normalnih dilatacija podu etiri vertikale na rastojanju tri etvrtine radijusa ipa od ose ipa. Uvodei u razmatranje modul elastinosti betona normalni naponi u pravcu ose su procenjivani kao i naponi ispod noice ipa. Procena je obavljena usvajajui konstantnu vrednost prenika ipa. Izmerene vrednosti su uporeene sa rezultatima numerikih simulacija ponaanja ipa i tla za sve sukcesivne faze optereivanja. Merenje dilatacija izvan tela ipa tokom testa pod vertikalnim i horizontalnim optereenjem u ovom sluaju nije imalo za cilj razvoj alternativne metode probnog optereenja ipa. To daje, na prvom mestu mogunost da se poblie ima uvid u dilatacije i napone najtee pristupanih delova razliitih tipova elemenata betonske konstrukcije, naroito ipova i drugih tipova dubokog fundiranja. Predstavljen aje tehnologija mlnitoringa pokazala se da koa veoma tana i konsistentna. KLJUNE REI: ipovi, duboki temelji, statiko probno optereenje, tehnologija merenja dilatacija,

elasto-plastino modeliranje, metoda konanih elemenata INTRODUCTION The bearing capacity and settlement of a vertically loaded pile can be estimated using a variety of methods (Bowles, 1996 and krabl, 2002), as can the passive earth resistance of the retaining walls (Vrecl-Kojc, 2005, krabl and Macuh, 2005, and Vrecl-Kojc and krabl, 2007). In order to identify the piles behaviour and its interaction with the surrounding soil layers, knowledge of the state of the strains along the pile axis inside the pile body is of great importance (krabl, 2008). To make this possible, a chain of measurement points has to be included outside the pile structure, and this chain should be fixed to the appropriate position on the pile reinforcement before the reinforcement is placed in the pile pit. After that the standard procedure for concreting the pile must be executed. The first opportunity for testing this idea was an estimation of the testing pile shafts resistance, as described in the papers of trukelj et al. (2005 and 2009). The basis of this estimation was a measurement of the normal strains of the pile in its axial direction at measuring points, distributed over equal distances along the pile axis. These strains are proportional to the axial forces in the pile. When the course of the axial force along the pile axis is known, the resistance of the pile shaft can be estimated. The first five measurement points in each of the four measurement chains were distributed over equal distances of 1.00 m, starting 0.75 m above the pile toe. The remaining three measurement points in each measurement chain were distributed over distances of 1.50 m. Since the strain gauges, the electrical contacts and the communication cables are very sensitive, any moisture and mechanical loading could be very harmful to their performance. Therefore, the measuring points were protected with special care. The measurement system fulfilled expectations and the measured results were accurate and stable. The only disadvantage of this measurement system is its non-flexibility. It can only be used to build measurement chains where the measurement points are placed along one line and oriented in the same direction. The loading test included two loading cases, i.e., the vertical and the horizontal. Each represented the introduction of the force in predefined loading steps: vertical in steps of 250 kN and horizontal in steps of 100 kN. The location of the testing pile was chosen to be on a construction site of the covered excavation 8-1 on the under section of the highway PonikveHrastje, Slovenia.The geological conditions of the testing piles location and the positions of the 32 sensors are shown in Fig.1, and described in Section 2.1. In the following sections, the measuring equipment, its installation, the performance of the static vertical and horizontal loading test, and, finally, the evaluation of the measurement results is presented. In the last part of the paper a comparison of the measured results with the results of numerical axis-symmetry and three-dimensional analyses using the finite-element method (FEM) are presented.

27

PRELIMINARY WORK AT THE TESTING SITE Geological field and laboratory investigations The geological conditions of the wider location of the testing site were acquired from the geotechnical report for the design of a covered excavation on the highway deviation (GI ZRMK, 2006). The strength and deformability parameters were defined on the basis of field investigations using a standard penetration test and probe measurements, as well as by laboratory testing of samples taken by sounds of depths up to 20 m from seven different locations. The half-space of the region of the covered excavation comprises clayey and clayey-gravel layers of depths 9.0 to 15.0 that lie on a limestone that is weathered in the first 13 m. At the location of the testing pile, additional sounding works during the pile pit excavation to a depth of 10.0 m were performed. The strength and deformability parameters at some depths of this additional sounding works on the basis of field investigations by standard dynamic penetration tests, and also on the basis of the laboratory testing of samples, were defined. The ground water level in this sounding was not encountered. The cross-section of the ground space with the pile (Fig. 1.) is composed of an original space with two characteristic layers: a 12-m clayey gravel overlays limestone weathered and compact. Fig. 1 shows the geological conditions of the testing piles location and the positions of the 32 sensors, ordered in four vertical measurement chains. Each vertical measurement chain was assigned a different colour in order to avoid confusion during the connection of the cables to the data-acquisition unit. Table 1 presents the strength parameters and the classification ( is the internal friction, c is the cohesion, Eoed is the oedometer elasticity modulus, and is unit weight) of the soil layers presented in Fig. 1, which were determined on the basis of laboratory and field testing on additional soil samples.

28

Figure 1. Section of the testing site with the disposition of the soil layers and the measuring points on the pile. Table 1. The strength parameters of the soil layers.

depth [m]

[]

c [kPa]

Eoed [kPa]

[kN/m3]

soil classification

0.0-12.0 32.0 28.0 10.000 19.0 GC/CH 12.0-14.0 40.0 3.000.0 100.000 25.0 weathered limestone

14.0 50.0 150.000.0 10.000.000 27.0 limestone Measuring equipment Besides the standard equipment that is needed for an estimation of the bearing capacity of the pile on the basis of a static loading test (ASTM, 1994), additional specially produced strain sensors were built into the pile body. They were placed along four verticals, circularly beside the longitudinal reinforcement bars of the pile (Fig. 1). The patented sensor design used for this purpose is very efficient, easy to build in, and robust enough to stand the water pressure and all the possible mechanical burdens during the concreting of the pile. Such a sensor can be placed in

29

the desired position in a very short time. It is insensitive to moisture and dust and its vital parts are very well protected against mechanical damage. The basis for such a sensor is a standard reinforcement bar of length about 150 cm and a diameter of 16 mm. In the middle of the reinforcement bar its surface is ground on one or both sides, depending on the number of strain gauges that should be installed. These strain gauges can be connected in a full, half, double-quarter or quarter Wheatstone bridge (Hoffman, 1989 and 1996). The connecting cable is protected by a polyethylene tube and fixed to the reinforcement bar. The protection coating consists of two layers of polyurethane varnish, a layer of special silicone putty and a layer of permanently plastic sealant putty coated with aluminium foil. This combination of protection was tested and remained waterproof even 30 m under water. The final layer represents the physical protection and can be made of polyethylene tube (Fig. 2) when the dimensions of the measured concrete elements are not too small to be significantly weakened by the built-in sensor. Otherwise, the physical protection of the measurement area can be made of cement mortar with the addition of an acrylic emulsion. This type of strain sensor proved to be very reliable, easy to build and cost effective, and can also be used together with the appropriate equipment and software (Brinkgrave and Vermeer, 1998 and 2005) for the purposes of monitoring the other parts of structures. The successive phases of the sensors preparation and their placing in the planned positions of the pile reinforcement are shown in Figs. 25.

Figure 2. The first phase of the sensor preparation: applying the strain gauges to the specially prepared

surfaces of the reinforcement bar and wiring.

Figure 3. Finished sensor after applying all the layers of protection against moisture and physical protection

Figure 4. The strain sensor fixed to the reinforcement of the testing pile

Figure 5. The pile reinforcement equipped with strain sensors prepared for placing into the pile pit.

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Construction of the testing pile The testing pile was constructed as a bored, reinforced concrete pile (Bowles, 1996) of 80 cm in diameter. The final length of the testing pile was 10.5 m. The length of the underground part of the pile was 10.0 m. As the thickness of the clayey gravel and gravel layer was about 12.0 m, the toe of the testing pile was about 2.0 m above the limestone base. After the pile was cleaned up to the level of the working area a 0.5-m-high reinforcement of the pile head was additionally concreted (Fig. 6). Installation of the static loading equipment The vertical and horizontal loads were applied with a hydraulic press. In the case of the vertical loading the hydraulic press was installed between the pile cap and a system of two mutually perpendicular girders positioned on the top of four columns, rigidly connected to the frame-shaped base surrounding the pile cap. Both steel girders were anchored to the limestone layer by four geotechnical anchors for the vertical load (Figs. 6 and 7). In the case of the horizontal loading the hydraulic press was installed between the vertical plane of the pile cap and the reinforced concrete block, which was leaned against the slope (Figs. 6 and 7).

Figure 6. The top of the testing pile before the

installation of the loading equipment. Figure 7. The testing site during the static load test.

THE PERFORMANCE OF THE STATIC LOADING TEST Static vertical loading test During the vertical test the load was applied by a hydraulic press placed between the steel girders and a horizontal surface of the pile-head at regular steps of 250 kN until the serviceability limit state of the pile (2500 kN) was reached. The serviceability limit was state established on the basis of the vertical loads, predicted by the design

31

project. After the first load step the pile was unloaded. Then it was re-loaded with the same intensity of load. All further load steps were applied without any interim relief. The next load step was performed only when the vertical displacement of the pile head, measured by the inductive displacement gauge simultaneously with the pile deformations, eased. The disposition of the measurement and loading equipment is shown in Fig. 8. Static horizontal loading test Before the horizontal test the soil was excavated to a depth of 2 m opposite the pushing side. The load was applied by a hydraulic press placed between the additional foundation and a vertical surface of the pile-head at regular steps of 150 kN until the serviceability limit state of the pile (750 kN) was reached. The serviceability limit was state established on the basis of horizontal loads, predicted by the design project. All the load steps were applied without interim relief. The next load step was performed only when the horizontal displacement of the pile head, measured by inductive displacement gauge simultaneously with pile deformations, eased.

Figure 8. Measurements during the vertical loading

test. Figure 9. The configuration of the loading and measurement equipment during the horizontal loading test.

EVALUATION OF THE MEASUREMENT RESULTS Evaluation of the measurement results for the vertical test The time course of the displacement measurements using the inductive displacement transducer during successive steps of the loading and unloading is shown in Fig. 10. In Fig. 11 the time course of the normal strains for one vertical measurement chain assigned a yellow colour (see Fig. 1) for each of eight levels is shown. The review of the peak values of the measured strains for the last vertical loading step at all measurement points of the same vertical measurement chain are shown in Fig. 12 as a curve representing the normal strains versus the distance along the pile axis. Since the strain values at the pile head and the pile toe could not be directly measured, they were extrapolated. In each of eight levels the average strains of all four measurement chains were calculated. The review of the peak values of the measured strains for the last vertical loading step at all the measurement points of the same vertical measurement chain are shown in Fig. 12 as a curve representing the normal strains versus the distance along the pile axis. Since the strain values at the pile head and the pile toe could not be measured directly, they were extrapolated. From the obtained average strains the pile contraction can be evaluated (Table 2). To evaluate the pile-toe settlement the obtained pile contraction should be subtracted from the measured value of pile heads vertical displacement. In our case the pile toe was 0.28 mm. For each of the eight levels the average strains of all four measurement chains were calculated. By taking the value of the Youngs modulus of the pile concrete to be E = 27.5 GPa, the curve of the normal stresses along the pile axis was obtained (Fig. 13). If the pilot is divided into segments whose boundaries are defined by the levels of the monitoring sites, on both ends of each segment the resulting axial force can be calculated from the average normal stress. Given that each of the segments are in balance, the force that balances the difference between the axial force

32

on the top of the segment and the axial force on the bottom segment thus represents the resultant shear stresses along the shaft of the segment. From these values the average values of the shear stresses on the pile shaft shown in Fig. 14 can be easily obtained. This is data from which it is possible to slim down the dimensions of the pilot in the event that the expected computational bearing-capacity values are distinctly smaller than the values obtained on the basis of the measurement results.

Table 2: The evaluation of the partial and cumulative pile contractions

distance from the

pile head [m]

average strains [m/m]

distances between levels [m]

middle strain values [m/m]

partial pile contractions [mm]

cumulative contractions [mm]

0.00 -218.15 0.75 -217.24 -0.163 -0.163 0.75 -216.33 1.50 -210.02 -0.315 -0.478 2.25 -203.70 1.50 -191.92 -0.288 -0.766 3.75 -180.14 1.00 -108.84 -0.109 -1.255 5.25 -147.24 1.00 -134.28 -0.134 -1.146 6.25 -121.31 1.50 -163.69 -0.246 -1.012 7.25 -96.36 1.00 -85.33 -0.085 -1.340 8.25 -74.30 1.00 -65.34 -0.065 -1.405 9.25 -56.37 0.75 -50.87 -0.038 -1.443* 10.00 -45.37 * The value of the entire pile contraction

33

Figure 10. Vertical displacements versus time measured by the inductive displacement transducer during successive vertical loading/unloading.

Figure 11. Measured strain signal recorded in the vertical chain of sensors assigned with a yellow colour for successive vertical loading/unloading.

34

-10

-9

-8

-7

-6

-5

-4

-3

-2

-1

0

-250.00 -200.00 -150.00 -100.00 -50.00 0.00

dist

ance

from

the

pile

hea

d [m

]

strain [mm/m]

Figure 12. The review of the peak values of the measured strains for the last vertical loading step at all the

measurement levels (the shown strain values represent the average of all four values measured at each level; the values at the pile head and pile toe are extrapolated)

-10

-9

-8

-7

-6

-5

-4

-3

-2

-1

0

-8.00 -6.00 -4.00 -2.00 0.00

dist

ance

from

the

pile

hea

d [m

]

average normal stress [MPa]

-10

-9

-8

-7

-6

-5

-4

-3

-2

-1

0

-250 -200 -150 -100 -50 0

dist

ance

from

the

pile

hea

d [m

]

average shear stress on the contact surface [kPa]

Figure 13. The review of the average values of the

normal stresses in the piles axial direction calculated from the averaged peak strain values for the last

loading step

Figure 14. Curves of the average values of the shear stresses on the pile shaft calculated from the average values of the normal stress for the last loading step

35

Evaluation of the measurement results for the horizontal test The time course of the displacement measurements using two inductive displacement transducers during successive steps of the loading and unloading for the horizontal test is shown in Fig. 15. The normal strains along the four measurement chains (yellow, green, blue and red in Fig. 1) of the pile were simultaneously measured. In Fig. 16 the time course of the normal strains for the measurement chain assigned with the green colour (see Fig. 1) for each of the eight levels is shown. The review of the peak values of the measured strains for the last vertical loading step at all the measurement points of the same vertical measurement chain are shown in Fig. 17 as a curve representing the normal strains versus the distance along the pile axis. Since the strain values at the pile head and the pile toe c