en o > (M SO O O (M n. štrukelj et al. inovrtive solution principles of wave problems in horizontrly layered media j. Klopčlč et al. displacements in the exploaatory tunnel ahead of the excavation face of šentvid tunnel o.žlender & l. Traunen the influence of poaosity on geomechanical chaaacteaistics of snail soil in the ljubljana marsh y. zhu et al. methods for control of seepage in rcc drms with watertight rnd drainage mersures ■H o ir> CO u> m fiCTfl GGOTGCHNICñ SLOVGNICñ Founders ISSN: 1854-0171 ustanovLteLjL Univerza v Mariboru, Fakulteta za gradbeništvo University of Maribor, Faculty of Civil Engineering ^Sll^ Univerza v Ljubljani, Fakulteta za gradbeništvo in geodezijo University of Ljubljana, Faculty of Civil and Geodetic Engineering . ; Univerza v Ljubljani, Naravoslovnotehniška fakulteta University of Ljubljana, Faculty of Natural Sciences and Engineering Slovensko geotehniško društvo Slovenian Geotechnical Society Društvo za podzemne in geotehniške konstrukcije Society for Underground and Geotechnical Constructions urednLškL odbor edLtorLaL Board LzdajateLj 1pubLLsher Univerza v Mariboru, Fakulteta za gradbeništvo University of Maribor, Faculty of Civil Engineering odgovornL urednLk edLtor-Ln-chLef Ludvik Trauner Univerza v Mariboru urednLka co-edLtors Stanislav Škrabl Univerza v Mariboru Bojan Žlender Univerza v Mariboru TehnLčna urednLka Desk edLtors Bojana Dolinar Univerza v Mariboru Borut Macuh Univerza v Mariboru LektorLča proof-Reader Metka Brkan NakLada cLrčuLatLon 500 izvodov - issues TLsk prLnt Tercia tisk d.o.o. Ptuj Revija redno izhaja dvakrat letno. Članki v reviji so recen-zirani s strani priznanih mednarodnih strokovnjakov. Baze podatkov v katerih je revija indeksirana: ICONDA - The international Construction database, GeoRef Pri financiranju revije sodeluje Javna agencija za raziskovalno dejavnost republike Slovenije. Darinka Battelino Università degli Studi di Trieste József Farkas Budapesti Müszaki és Gazdaságtudományi Egyetem Theodoros Hatzigogos Aristotle University of Thessaloniki Rolf Katzenbach Technische Universität Darmstadt Zlatko Langof Univerzitet u Sarajevu Jakob Likar Univerza v Ljubljani Janko Logar Univerza v Ljubljani Bojan Majes Univerza v Ljubljani Milan Maksimovic Univerzitet u Beogradu Borut Petkovšek Zavod za gradbeništvo Slovenije Mihael Ribičič Univerza v Ljubljani César Sagaseta Universidad de Cantabria Stephan Semprich Technische Universität Graz Abdul-Hamid Soubra Université de Nantes Ivan Vaniček Ceské vysoké učeni technické v Praze Franjo Veric Sveučilište u Zagrebu Address NasLov urednLštva ACTA GEOTECHNICA SLOVENICA Univerza v Mariboru, Fakulteta za gradbeništvo Smetanova ulica 17 2000 Maribor Slovenija Telefon / Telephone: +386 (0)2 22 94 300 Faks / Fax: +386 (0)2 25 24 179 E-pošta / E-mail: ags@uni-mb.s spLetnL nasLov http:IIwww. fg.uni-mb.siIjournal-ags I web Address The journal is published twice a year. Papers are peer reviewed by renowned international experts. Indexation data bases of the journal: ICONDA - The International Construction Database, GeoRef Financially supported also by Slovenian Research Agency. VSEBINA fr Ludvik Trauner uvodnik Andrej Štrukelj In drugI inovativni pristopi k reševanju valovnih problemov v horizontalno slojevitem polprostoru a jure Klopčič In drugI pomiki v raziskovalnem rovu pred izkopnim čelom predora šentvid Bojan Žlender in Ludvik Trauner vpliv poroznosti na geomehanske lastnosti polžarice iz ljubljanskega barja [44- Yueming zhu in drugi metode za kontrolo pronicanja vode skozi pregrade iz valjanega betona z meritvami vodotesnosti in drenaž navodila avtorjem CONTENTS Ludvik Trauner editorial Andrej Štrukelj et al. inovrtive solution principles of wave problems in horizontrly layered medirs jure Klopčič et al. misplacements in the exploratory tunnel ahead of the excavation face of šentvid tunnel E [35] Bojan ¿lender and Ludvlk Trauner the influence of porosity on geomechrnicrl characteristics of snail soil in the ljubljana mrrsh Yuemlng zhu et al. methods for control of seepage in rcc dams with watertight and drainage measuress [¿Ž] instructions for authors E UVODNIK Mednarodna revija Acta Geotechnica Slovenica, ki redno izhaja že tri leta, postaja vedno bolj odmevna tako doma kot v tujini, k čemer pripomore tudi dejstvo, da jo sedaj lahko nejdemo v različnih specializiranih bazah podatkov kot sta Iconda - The International Construction database in GeoRef. Druga številka tretjega letnika prinaša štiri zelo zanimive prispevke domačih in tujih avtorjev. V prvem članku A. Štrukelj, A. Umek in T. Pliberšek predstavljajo inženirsko smiselno transformacijo površinskih pomikov horizontalno slojevitega polprostora, kjer so prikazani tipi valov, ki nastopajo v polprostoru. Pokazano je, da so površinski valovi izraženi kot residumi izračunani v poljih integranda, volumski valovi pa so izraženi kot integrali vzdolž razvejiščnih rezov. Singularnost, ki se v osnovni singularni rešitvi v elastodinamiki pojavlja vedno, je eksaktno izločena. V drugem delu prispevka je detajlno obdelan problem pojava in obnašanja Stonely-jevih valov. Drugi članek J. Klopčiča J. Logarja, T. Ambrožiča, A. Štimulaka, A. Marjetiča, S. Bogatin in B. Majesa predstavlja projekt predora Šentvid ter metodo 3D meritev pomikov z rezultati teh meritev. Pridobljeni rezultati so interpretirani skladno z ugotovljeno geološko sestavo hribine s poudarkom na parametrih, ki so pomembni za bolj zanesljivo uporabo pomi-kovne funkcije. V tretjem prispevku avtorja B. Žlender in L. Trauner obravnavata mineraloške in fizikalne lastnosti melja (polžarice z Ljubljanskega barja) in njihov vpliv na vrednosti parametrov geomehanskih lastnosti. Rezultati raziskave kažejo na sovis-nost geomehanskih lastnosti in poroznosti. Odnose lahko izrazimo kot funkcije gostote oz. poroznosti ali vlažnosti. Iz rezultatov je razvidno, da so spremembe vodoprepustnosti, konsolidacije in stisljivosti nelinearno odvisne od sprememb poroznosti. Spremembe mehanskih parametrov kot so elastični modul, Poissonov količnik in strižni kot, so pri manjših spremembah poroznosti neizrazite in skoraj linearne. Avtorji J. Zhu, S. Semprich, E. Bauer, C. Yuan in D. Sun v četrtem prispevku podajajo teoretične osnove ter rezultate terenskih raziskav o pronicanju vode skozi valjani beton, ukrepih za doseganje vodotesnosti, drenaži, optimalnem načrtovanju in nadzoru pronicanja. Rezultati prikazane študije so bili koristno uporabljeni pri izgradnji in povratnih analizah številnih jezov. Članek podrobno opisuje raziskovalne dosežke študije, teoretske osnove ter njihovo praktično uporabo pri visokih gravitacijskih jezovih iz valjanega betona. Ludvik Trauner Glavni urednik EDITORIAL The international journal Acta Geotechnica Slovenica, which has now been published for three years, is acquiring increasing national and worldwide reputation. An indication of its success may also be seen the fact that the journal is being abstracted in different specialized databases such as Iconda, the International Construction Database, and GeoRef. The second issue of Year 3 contains four interesting articles authored by home and foreign contributors. A. Štrukelj, A. Umek, and T. Pliberšek describe an engineeringly reasonable transformation of surface displacements of a horizontally layered half-space, where different types of waves are present. It is shown that surface waves are expressed through residues in poles of the integrand and the volume waves are expressed as integrals along corresponding branch cuts. The singularity which always appears in the basic singular solution in elastodynamics is explicitly excluded in this case. In the second part of this article, the appearance and behaviour of Stonely waves are investigated in greater detail. The second article by J. Klopčič, J. Logar, T. Ambrožič, A. Štimulak, A. Marjetic, S Bogatin, and B. Majes presents the Šentvid tunnel project, the method of 3D displacement measurements, the results of these measurements and their interpretation according to the geological structure of the site. The emphasis is placed on items which play an important role for a more efficient application of the displacement function. B. Žlender and L. Trauner describe mineralogical and physical characteristics of snail soil in the Ljubljana marsh and their influence on parameter values of geomechanical characteristics. The results of the tests show that interdependency exists between geomechanical characteristics and porosity. These relationships can be expressed as functions of density, porosity or water content. It is evident from the results that changes of the coefficient of permeability, the coefficient of consolidation, and the coefficient of volume compressibility are non-linear with respect to changes of porosity. Changes of mechanical parameters, such as Young modulus, Poisson ratio, and friction angle are indistinct and almost linear at lower changes of porosity. J. Zhu, S. Semprich, E. Bauer, C. Yuan, and D. Sun describe theoretical bases and the results of field tests performed on roller-compacted concrete tested for water seepage. Water-tightness measures, drainage and optimal design and control of seepage are presented. The results of this study have been successfully applied in construction and in back analyses of several dams. The article describes the research findings in detail, focusing on theoretical bases and their applications in high roller-compacted gravity dams. Ludvik Trauner Editor-in-chief INOVRTIVNI PRISTOPI K REŠEVANJU VALOVNIH PROBLEMOV V HORIZONTALNO SLOJEVITEM POLPROSTORU andrej štrukelj, andrej umek in tomaž pliberšek o avtorjih Andrej Štrukelj Univerza v Mariboru, Fakulteta za gradbeništvo Smetanova ulica 17, 2000 Maribor, Slovenija E-pošta: andrej.strukelj@uni-mb.si Andrej Umek Univerza v Mariboru, Fakulteta za gradbeništvo Smetanova ulica 17, 2000 Maribor, Slovenija E-pošta: umek@uni-mb.si Tomaž Pliberšek Univerza v Mariboru, Fakulteta za gradbeništvo Smetanova ulica 17, 2000 Maribor, Slovenija E-pošta: tomaz.plibersek@uni-mb.si Izvleček V članku je predstavljena inženirsko smiselna transformacija površinskih pomikov horizontalno slojevitega polprostora, kjer so prikazani tipi valov, ki nastopajo v polprostoru. Pokazano je, da so površinski valovi izraženi kot residui izračunani v polih integranda, volumski valovi pa so izraženi kot integrali vzdolž razvejiščnih rezov. Singularnost, ki se v osnovni singu-larni rešitvi v elastodinamiki pojavlja vedno, je eksaktno izločena. V drugem delu prispevka je detejlno obdelan problem pojava in obnašanja Stonely-jevih valov. Pokazano je, da v primeru slojev končne debeline njihov pojav in hitrosti ne zavi-sijo le od materialnih karakteristik sosednjih slojev, ampak tudi od njihove debeline. Ključne besede horizontalno slojeviti polprostor, volumski valovi, površinski valovi, Stonely-jevi valovi, Green-ova funkcija INOVATIVE SOLUTION PRINCIPLES OF WAVE PROBLEMS IN HORIZONTALY LAYERED MEDIAS andrej štrukelj, andrej umek and tomaž pliberšek About the authors Andrej Štrukelj University of Maribor, Faculty of Civil Engineering Smetanova ulica 17, 2000 Maribor, Slovenia E-mail: andrej.strukelj@uni-mb.si Andrej Umek University of Maribor, Faculty of Civil Engineering Smetanova ulica 17, 2000 Maribor, Slovenia E-mail: umek@uni-mb.si Tomaž Pliberšek University of Maribor, Faculty of Civil Engineering Smetanova ulica 17, 2000 Maribor, Slovenia E-mail: tomaz.plibersek@uni-mb.si Abstract The paper represents engineeringly reasonable transformation of surface displacements of horizontaly layered half-space. The latter shows in the half-space present types of waves. It is shown that surface waves are expressed through residues in poles of the integrand and the volume waves are expressed as integrals along corresponding branch cuts. The singularity which always appears in the basic singular solution in elastodynamics is in this case exactly excluded. In the second part of the paper the appearance and behaviour of Stonely waves is investigated in greater detail. It is shown that in the case of layers of finite thickness their appearance and velocities depends not only on the material characteristics of neighbouring layers but also on their thickness. Keywords horizontaly layered half-space, volume waves, surface waves, Stonely waves, Green's function 1 INTRODUCTION Waves, which are generated in the source of vibrations due to natural (f. ex. earthquakes), technical (f. ex. railway traffic, pilot driving) and man-made agitations, are propagated through the soil and disturb the functioning of sensitive instruments, cause human discomfort, and can possibly lead to structural damage. Investigating wave propagation is one of the main interests of applied mechanics and civil engineering because of its importance for dynamic structure-soil interaction, foundation engineering, seismology, and geophysical methods for determining the structure of foundation soils. As a rule, the soil is modeled as a half space in all problems displaying local soil characteristics or soil movement in the immediate vicinity of the selected point. The problematic nature of soil propagation in three-dimensional homogeneous and mainly layered half-space mainly has aroused significant scientific attention. Knowledge of soil propagation in a layered half-space is thus of key importance and has always attracted much scientific attention. Bromwich [1] was the first to investigate wave propagation in half-continuous solid medium covered with a solid layer of continuous thickness. His work, which refers to standing waves with wave lengths greater than the thickness of the top layer, was continued by Love [2] who also investigated waves with equal or smaller lengths than the thickness of the top layer. Sezawa [3] studied the dispersion of elastic waves which propagate on the surface of layered bodies and on curved surfaces, whilst Thomson [4] investigated the transmission of elastic waves through a layered solid medium. Dispersion of surface waves in a multi-layered medium was studies by Haskell [5]. The classical work by Ewing, Jardetzky and Press [6], which also summarized contributions of several other authors, must also be mentioned in this review. The authors of this contribution have chosen Green's function as a starting point for wave propagation in horizontally layered half-space. They [7] first formulated Green's function for the elastic layer loaded on the surface with concentrated harmonic force of a general direction. Then, by taking into account the findings of Kobayashi [8], they expressed Green's function of a halfspace, i. e. a layer of continuous thickness described with integrals for a half-continuous integration field, with integrals along suitably chosen branch cuts and residues in the poles of integrands. The solutions described in this article lead to the exact solution of Green's function of a horizontally layered half-space described in the article [9]. It results from the above quoted works that it is possible to express the displacements on the surface of a horizontally layered half-space as a sum of products of integrals and corner functions due to a concentrated time-dependent force of general direction acting on the surface. The concentrated force of general direction can namely be presented with two components, i. e. a vertical and a horizontal one. The fact that the solution of integrals using discreet Fourier-Bessel transformation does not provide satisfactory accuracy and convergence near the source point especially motivated the authors to adopt an alternative three-step approach to the solution of the integrals. As shown in this article, after the adoption of three steps, the integrals in the form that does not observe the singularity where it exists are written as integrals along branch cuts and residues in the poles of the integrand. This confirms the facts originating in the theory of elastodynamics, namely that it is possible to express surface waves in an elastic medium with poles and that volume waves are given with integrals along nodal cuts. This approach has opened completely new possibilities to study surface waves. This article presents the results of the new approach. In passing from a homogenous half-space to a horizontally layered half-space, Stonely waves can be generated besides Rayleigh waves. The magnitude of the system of equations for determining integration constants quickly increases with the increase of the number of layers, which is evident from the article by [9]. This also means that Stonely poles can occur beside Rayleigh poles in solutions. It is thus reasonable to investigate the phenomenon of singularity in the integrand more in detail. In the continuation, the article will show at which ratios of thickness and shear modulus of two adjoining layers Stonely waves occur and how the thickness of layers affects the speed of Stonely and Rayleigh waves. 2 THREE STEP SOLUTION The thesis that it is engineeringly reasonable to transform surface displacements of a layered half-space into the from which is easy to calculate and which clearly shows wave types present in the half-space, is emphasized in the works [10], [11], [9]. For an axis-symmetrical example of vertically concentrated load [9] exerted on the surface of a top layer, as well as for a tangentially loaded layered half-space [11], displacements are expressed as a sum of integrals: I (r) = f F (r,)-J n (n ■ r )■ dV (1) 0 multiplied with adequate trigonometric functions with a circumferential coordinate as an argument. In these integrals, J is the Bessel function of the first kind of order and n is the order of the function which can adopt the values of 0, 1 and 2. It is known from scientific literature that the evaluation of these integrals with discrete Fourier-Bessel transformation does not lead to required accuracy and stability, especially when the integrals become singular at r ^ 0 . Mathematical analysis and mechanical understanding of the problem result in an alternative approach to the evaluation of the integrals shown in [11]. The authors evaluated the integrals of type (1) in three steps using the so-called method of the »expected shape«. The first step is the extraction of singularity. In an innovative way and observing the fact that lim I(r) = lim I F(rj)- J0 (r ■ r)- dri- rrJ -- I lim F (r )■ J0 (r ■ r )■ dr = C ■ IJ0 (r ■ r )■ dr = — J r^œ J r (2) the integrals which contain singularity are divided into two parts: a regular integral and a singular integral: I(r) = -■ C + Il (r) , (2a) where I1 (r) is: I1 (r) = J\f(r)-c ■ Jo(r ■ r)■ dr. (3) A singular integral has a simple integrand so that it can be analytically solved, whilst regular integrals are better to be calculated using discrete Fourier-Bessel transformation. The next step leads to their even simpler 0 0 0 calculation. It is namely evident that the below functions are present in the function F(r) of the integrand: a =Vn2 - y2 , ßi = Vn2 - $ , e±a ■h, e ±ßi h (4) where Y, = hiL ; ^ = (4a) kLi or kTi, respectively, are wave numbers of longitudinal or transversal waves, and h, is the thickness of the i-th layer. So, it is reasonable to substitute the mechanical understanding of the function F (rj) with its analytical approach and a suitable selection of branch cuts to make the function uniform. A branch cut, which is shown in Fig. (1), was first selected inventively. If we wish that the function F(rj) is an even function of the variable r/ on the real r/ -axis we must modify accordingly exponential functions (4) which are neither even nor odd. To make these exponential functions uniform and even they are replaced with their analytical continuation: ±ri a h e hi = e Inl e±ß' h,= e n . The mentioned modification, which makes the integrands even, allows the Bessel function to be divided into two parts: 2 J, (z) = h (z) + h (-z) (6) and consequently to transforms Hankel's inverse integrands, which are present in the derived components of Green's function, into integrals with the integral range from minus infinite to infinite: 11 (r) = ^[f(n)-C]• J (n- r) •dr 0 1 fTO[ ] 0 [ ] 1 = --|/[f(r)-C| •hi (r -r) •dr + J|f(r)-^ •h, (r -r)• dr\ (7) 1 r i i = - • [If (r)-cl-h, (r • r)-dr = - • i2 (r) 2 -M J 2 The final or the third step is the evaluation of the integral 12 (r). It is shown in [11] that this integral can be evaluated with a contour integration in the complex plane of the Hankel parameter. Based on the theorem of residua [12] we obtain: I2+ + I2R + I2- + I2b1 + I2r + I2b2 = 2n 1 ^reS , (8) where the direction of the integration is evident from Fig. 2. (5) material damping, which is expressed with the complex shear modulus fj,i = ßoi ■ erf , is considered. Figure 2. The integration path of the evaluation of Hankel inverse integrals which are present in the components of Green's function. The first member of the c indexes of the expressions in this figure is valid for all displacements of a horizontally layered half-space surface. Material damping is considered for clearer presentation. Expression (8) is reasonably re-arranged: I2 = I2- + I2+ = 2 n i ^ res - I2b1 - I2r - I2b2 , (9) or written in the form: I = - + 2n i ^ fes - I2b\ - I2r - 12b2 . (10) Inverse Hankel integrals, which occur in the components of Green's function, adopt the below final form using the described alternative three-step approach: C I = r + 2 n i ^ res + ?(,)•J„ (n • r )• dr, . (11) Equation (11) provides engineeringly reasonable transformation of displacements of a layered half-space surface because it transparently shows the types of waves present in the half-space. The first member presents singularity which always appears in the basic singular solution in elastodynamics and which is excluded from inverse Henkel integrals that define the components of Green's function. Linking the findings of mathematical physics with the findings in theoretic and applied mechanics leads to the conclusion that surface waves in Eq. (11) are expressed with residues in the poles of the integrand, whilst volume waves are expressed with integrals along branch cuts. 3 STONeLY wnves In evaluating Green's function it was shown that the contribution of the second member of Eq. (11) is very important. As already said, it manifests the contribution of surface waves which is defined with the poles of integrands. In a homogenous half-space, two conjugated complex poles only appear which define the presence of Rayleigh waves. The system of equations quickly increases with the increase of the number of layers, due to the requirement that continuity conditions on contact planes of individual layers be fulfilled. So, the equation system matrix for determining unknown integration constants increases in accordance with equation 4 • n + 2 in case of vertical concentrated force acting on the surface. In case of tangential concentrated surface load, the size of the system matrix is dictated with equation 6 • n + 3 . In both cases, n presents the number of layers. Besides Rayleigh waves, which always appear on the surface of a half-space, Stonely waves can also appear on contact surfaces. The presence of surface waves is defined in the system of equations with zeros of the system determinant or with singularities of integrands in integrals for the inverse transformation of expressions for individual components of Green's function. The search for the integrand singularity becomes more demanding when the system increases; yet, numerical calculation of singularity greatly reduces the speed of evaluating Green's function and hinders the automation of the calculation process. Therefore, it is sensible to investigate the process of singularity in integrands more in detail. a Rayleigh's waves always appear when the medium through which waving is propagated has an unobstructed surface. Stonely waves appear on contact surfaces between individual layers, yet only at certain ratios of thicknesses and shear modula of two neighbouring layers. The range of the appearance of Stonely waves is investigated on an example of two half-spaces. Figure 3. Model of two half-spaces. As shown in Fig. 3, we take two homogenous elastic isotropic half-spaces with different thicknesses and different shear modula which are loaded at their contact surface with a vertical concentrated force. The problem is mathematically formulated so that the vector of displacement for each half-space is written by using potentials [13] in the form: u = V ■ p + Vx$ (12) The system of so linked partial differential equations, which presents the equations of movement, _ / _ d2U n ■ V2U + (A + V- (•U) = p—- (13) 1 ; dt disintegrates into the system of non-linked differential differential equations 2 1 d2^ =7'W -V= (14) (15) where cL and cT are the speeds of longitudinal or shear wave front, respectively. They are translated into the frequency domain with the Fourier exponential transfor- mation t ^ u . The studied elastodynamic problem is axially symmetrical, which dictates the use of the cylindrical coordinate system. Furthermore, the component of the displacement in the direction § equals zero, both potentials, the scalar one ty and the vector one ^ , must be independent of § , and therefore the components of the vector potential ^ in the directions r and z must also equal zero. The individual components of the vector of displacement therefore resume the form: dr dz 0 dyr + 1 d(r• V dz r dr (16) and partial differential equations (14) and (15): / \ 2 d2tyr 1 dtyr d2tyr dr2 r dr dz2 u V = 0 (17) d2^ ,1 , d2^ Vv dr2 r dr dz2 •Vv = 0 . (18) The relationship between the components of tension and displacements has the form: - T dw \ -z = 2 • ß---\ • dz _ du T zr = ß^ T.--ß dz u Si u Vr (19) (20) Partial differential equations (17) and (18) are transformed into ordinary differential equations using Hankel integral transformation r ^ £: dV (j) dz2 d2V (j) dz2 j2 - j2 - V (j ) = 0 (21) V1 (j) = 0 (22) The transformed components of the vector of displacement have the form w1 (£) = -£■ p (£)-d? (£) (23) dz 2 2 2 2 i 2 T w0 (i) = d00 (i) + i • ^ (i) (24) dz We then also transform the expressions for normal and shear tensions: i = kn • n (33) CT kL cL kT \X + 2• ß \ 2•( 1 — v) 1-2 • v ■ = 7 (34) U (i ) = ß •| 2 • i2 — u •0° (i) + 2 • i •dtf (i) (25) dz u1 (i)=— ß -|2 •i •dZ 00 (i) + 2 • i2 — ^(i)i• (26) ^2 = (35) General solutions for individual potentials therefore are: = Ci,i • exp(2 • n ' Vn2 -7i2 • C) = Cu •exp(2 •n• Ô! • C) (36) Transformed differential Eqs. (21) and (22) for the potentials tp° (£) and (£) are ordinary homogeneous differential equations where z is the only variable. Their general solutions are: 00 =f xxX; xxxxxxxj^. XXXxxx xxXv X measured displacements déplacement function For top heading excavation displacement function For bench excavation -O- top heading -D- bonch t invert . -0- bench * insert (left) (right) 1 Construction phases > 1 1 Date program Predor - V3.076 Figure 4. Displacement history plot of the vertical displacement of crown point with fitted displacement function for top heading and bench excavation. For the case described in Fig. 2 the fitted displacement function to the measured displacements is shown in Fig. 4. The divergence of the theoretical curve and measured displacements in this case indicated the support failure in the vicinity of the measuring section (Fig. 3). According to the fitted displacement curve the failure of the cross section occurred soon after the ring closure at approximately 30 cm of vertical displacement of the crown point. At this point additional support measures should have been undertaken to avoid further disintegration of the rock mass - support system. 2.3 limitations of the displacement function While part 3 of the displacement function (after the excavation face passed the observed cross section - Eq. 3) is very well defined by fitting the curve to the measured displacements, the first two parts (ahead of the face (part 1) and between the excavation and the support installation (part 2)) are determined on the basis of numerical simulations and rely on the parameters that depend on the rock mass behaviour ahead of the tunnel face: Q: = proportion of the total stress change due to the tunnel excavation that occurs ahead of the face and xf = parameter presenting the magnitude of the pre-face domain. The magnitude of both parameters and the shape of first two parts of the displacement function could be verified or accurately determined only by measuring the displacements within the rock mass ahead of the tunnel face. Obviously, this cannot be done in terms of ordinary geodetic methods. In case of low overburden and provided that monitoring of the displacements is established in the tunnel as well as on the surface along and perpendicular to the tunnel axis, the amount of the displacements that occurred before the first measurement in the observed cross section can be estimated from the difference between the tunnel and the surface settlements. Extensive monitoring of the surface displacements was established above the eastern part of the Trojane tunnel that passes the Trojane village at shallow depth. The Trojane tunnel was constructed in heavily faulted and densely foliated rock mass of carboniferous age. The comparison of the measured displacements in the tunnel tube and above it revealed that less than 50% of the final displacement measured on the surface was only measured in the tunnel tube [11]. The rest of displacement occurred before the first measurement of the measuring cross section was performed. From the relation between the surface displacement and the distance to the approaching top heading excavation face the magnitude of the pre-face domain can also be estimated. More reliable information on the rock mass behaviour ahead of the face and magnitude of the pre-face domain was given by the interpretation of the measurements of the horizontal inclinometer that was installed in the length of 40 m in the crown of the left tube of the Trojane tunnel under the Trojane village ([12], [13]). The crown settlements measured by the horizontal inclinometer are shown in Fig. 5 together with influence lines. At the chainage km 80+293 a geodetic measuring cross section (marked in Fig. 5) was installed. The settlement measured by the horizontal inclinometer at the given chainage was 17.5 cm and the measured spatial displacement of the crown point in the geodetic cross section was 4.9 cm, which is less than 30 % of the final settlement. All three components of the displacement vector are drawn in Fig. 6. The magnitude of the pre-face domain is approximately 1 - 1.5 equivalent tunnel diameter (10 - 15 m in the case of the Trojane tunnel) ahead of the face as one can observe from Fig. 5. The advantages of measuring displacements ahead of the face with the horizontal inclinometer are the accuracy of measurements, minimum disturbance of the rock mass and the position of the inclinometer, which allows the direct comparison with the geodetically measured settlement of the crown point. The major disadvantage is the inability of measuring the 3D displacement. Only the settlement of the crown is obtained. The longitudinal component of pre-displacement might be additionally measured using deformeters, but no reports on such measurements can be presently found in literature. 3D displacements ahead of the tunnel face can be easily obtained by geodetic monitoring, when a small diameter tunnel existed within the alignment of the future tunnel with considerably larger cross section. chainage km 80+xxx,xx 263,00 285,00 287,00 289,00 291,00 293,00 295,00 297,00 299,00 301,00 303,00 •Top Heading at 283,50 ■Top Heading Invert at 291.60 ■ Top Heading Invert to 296,71 Phase 4 to 299,30 Top Heading at 300.20 Top Heading invert to 287,68 - pipe roof {42 pipes) ■ Top Heading Invert to 297,51 • Top Heading invert at 299,30 ■Face position • Top Heading Invert to 289,41 - Top Heading Invert to 293.23 Top Heading Invert to 298,21 -pipe roof {42 pipes! ■ Trend -1.60 m ■ Top Heading Invert to 290,96 - Top Heading Invert to 295,20 - Top Heading Invert to 298,93 ■8ench to 287.50 Trend -3.20 m Figure 5. The influence lines of the settlements, measured by the horizontal inclinometer in the crown of the left tube of the Trojane tunnel (unpublished report by Button and Volkmann). Date Figure 6. Displacement history plot of all three components of the crown point displacement vector (left tube of the Trojane tunnel, measuring cross section at the chainage km 80+293). 3 3D GEODETIC DISPLACEMENT MEASUREMENTS AHAED OF THE MAIN TUNNEL FACE In the eight-month period between September 2005 and April 2006 the 3D displacement measurements were performed in the exploratory tunnel of the Šentvid motorway tunnel ahead of the main tunnel excavation face. 3.1 the šentvid tunnel The Šentvid tunnel system is the missing link of the Slovenian A2 Karavanke-Ljubljana motorway to the Ljubljana ring motorway (Fig. 7). Figure 8. The scheme of the Šentvid tunnel [14]. Figure 7. Slovenian road map. An approximately 1100 m long motorway tunnel is designed as a double tube tunnel with two large merging caverns with maximum excavation cross section of approximately 330 square meters and the length of 60 m (label A in Fig. 8). The Šentvid tunnel consists of twin two lane tunnels from northern portal up to the merging caverns (label B) and twin three lane tunnels from southern portal to the merging caverns (label C). Two ramp tunnels (label D) will connect the Celovška street to the main motorway tunnel. The tunnel system is currently under construction. All underground structures are constructed following the principles of NATM. Maximum overburden reaches 115 m. The Šentvid tunnel alignment passes through densely foliated clastic sedimentary rocks of carboniferous age, mainly sandstones, siltstones and clayey slates. The region has undergone intense tectonic deformations, presumably during several deformation phases. Due to intensive tectonics the rock is folded, fault zones are up to several meters thick and consist mainly of gouge clay. The rock mass itself is very heterogeneous and anisotropic (Fig. 9). The quantity of water that percolates from the surface into the tunnel tube is small. Water appears mainly in fault zones. Together with deformations that occur due to tunneling this water causes the increase of the water content along the foliation and consequently the decrease of rock suction, which affects mechanical properties of the rock mass and worsens the tunneling conditions in the vicinity of fault zones. Tunneling conditions for the Šentvid tunnel system were estimated in the range from fair to very poor [15]. Figure 9. Geological cross section of the Šentvid hill - left tube, northern part (ELEA iC). |M> v- left tube jr rinhf. h. cavern / cavern 4,20 ^ y tube ■ —r—j — 1 T O-- Figure 10. Ground plan and cross section of the Šentvid exploratory tunnel. To determine the most favorable position of the caverns in terms of geological and geotechnical criteria the exploration gallery in the axis of the main tunnel was constructed in the final stage of the design. The length of the exploration gallery was approximately 655 m with a cross section of 13 m2 (Fig. 10). Due to a small cross section the primary lining was executed as a micro fiber reinforced shotcrete with steel arches and wire mesh in the crown. No rock bolts were installed. The exploratory tunnel allowed the establishment of a reliable geological model and enabled the in-situ geotechnical testing (core drilling, geophysical surveys, extensometers). The geodetic measurements of the 3D displacements during the exploratory tunnel construction improved the knowledge of the rock mass behaviour and its response to the tunnel excavation. Beside all the information that contributed to the successful execution of the main tunnel, the exploration gallery enabled the observations of the rock mass - support system behaviour ahead of the face of the main tunnel during the execution of the main tunnel. 3.2 performance of geodetic measurements Continuous measurements of the primary lining were performed in both tubes of the exploratory tunnel. When the left tube of the main tunnel was approaching the intersection of the cross-passage and left tube of the exploratory tunnel (chainage km 1.3+97), the geodetic measurements of the targets were performed once a day in cross-passage, perpendicular to the main tunnel axis. These results improved the knowledge on the rock mass behaviour perpendicular to the tunnel axis. The monitored sections are marked with a blue dotted line in Fig. 10. The lengths of the monitored sections and the number of the measuring profiles are given in Table 1. Table 1. Chainages, lengths and number of measuring profilein monitored tunnel sections. chainage section length [m] measuring profiles right tube km 1.2+57 - km 1.3+32 75 37 cross-passage km 0+20 - km 0+40 20 9 left tube km 1.3+97 - km 1.5+44 147 72 The geodetic prisms were installed in the primary lining of the exploratory tunnel every two meters in the crown and every six meters on both side walls and ground along the tunnel axis. The 3D positions of the targets in the lining of the exploratory tunnel in front of the main tunnel face were recorded every hour by precise electronic tachymeter LEICA TCRP 1201R300. 3.3 results of 3d displacement measurements The magnitude of the measured displacements was in range from a few millimeters to almost 35 cm. The maximum observed displacements in each direction are given in Table 2. Figure 11a. Electronic tachymeter and geodetic prisms installed in the primary lining and on the bottom of the exploration gallery. Figure 11b. Position of the exploratory tunnel on the double lane tunnel top heading face. Table 2. Maximum displacements in each direction for different points in the exploratory tunnel due to the excavation of the main tunnel. vertical lateral longitudinal [cm]_dir:_[cm]_dir:_[cm]_dir. crown 18:7 I 6:4 —— 29:3 / left side wall 26:5 I 16:9 —>• 21:1 / right side wall 20:4 I 17:1 — 21:0 / bottom 34:6 j 11.3 — 6.4 / 3.3.1 General behaviour General behaviour of rock mass - support system that was measured in the exploratory tunnel in front of the main tunnel face is shown in Fig. 12 together with the prevailing geological scheme in monitored sections (sub-horizontal foliation in cross section with slight inclination towards left side of the tunnel and steeply inclined foliation into the excavation direction in longitudinal cross section, i.e. dip angle of the foliation is approximately 55° with relative dip direction with respect to the tunnel axis of 25° to the left). The distinctive bilinear deformation pattern can be observed in Fig. 12 (evident in longitudinal section and cross section). The displacement vectors in cross section of the exploratory tunnel followed the direction of the rock mass foliation towards the left sidewall of the tunnel when the excavation face was far away from the observed cross section. In this first deformation phase the additional rock mass pressure onto the primary lining was small and the sliding mechanism along the foliation dominated over the radial deformation due to rock mass pressure. For this reason the left sidewall point tended vertically down and not in the radial direction. Similar explanation can be given for the longitudinal section. As the tunnel face approached the observed cross section, the displacements due to the rock mass pressure became significantly larger than the sliding displacements along the foliation and consequently the displacement vectors changed their orientation. Figure 12. Bilinear displacement vectors in cross section and in longitudinal section in the left tube at chainage km 1.4+56. The deformation mechanism changed from sliding along the foliation in the first phase to the buckling of the foliation in the second phase when the influence of the excavation face to the monitored cross section was intensified. On the basis of several monitored cross sections with similar deformation patterns we can conclude that the rock mass - support system behaviour well ahead of the face is mainly governed by the orientation of the rock mass discontinuities. The displacement vector orientation change occurred suddenly and a turn point can be located on the plots of displacement vectors (Fig. 12). This turn point happened at a certain distance between the monitored section and the approaching face of the top heading. If the percentage of the final measured displacement is plotted against the distance from the top heading excavation face the displacement-distance curve is obtained (Fig. 13). This curve is also bilinear and can be simply divided in two parts. The turn point for monitored cross section between both parts occurred 5.5 m ahead of the two lane tunnel excavation face and at 45 % of the final displacement for the presented measuring cross section. The analysis of other cross sections indicates that the turn point occurs at about one half of an equivalent tunnel diameter in front of top heading face at 40-60% of final displacement. Between the measuring section and the approaching face of the top heading at the moment when the diagram in Fig. 13 exhibits the turn point there is the area of large displacements. The area further away from the tunnel face in the excavation direction can be referred to as the area of small displacements. Observed structural damage of the primary lining coincided well with the measured displacements. In the area of small displacements only micro cracks were noticed while in the section closer to the tunnel face up to 25 cm wide cracks were registered (Fig. 14). These large cracks were formed mainly due to large longitudinal displacements. As it can be seen from displacement vectors in longitudinal section in Fig. 12, the longitudinal displacements in the first phase were much smaller than the vertical, but were greatly increased in the second phase. In some cross sections no vertical displacements were observed in the second deformation phase, sometimes even heaving was registered. Figure 13. Relative vertical displacement of the crown point in exploratory gallery vs. distance to the top heading excavation face. Figure 14. Large crack in primary lining of the exploration gallery due to large longitudinal displacements. 3.3.2 Area of small displacements ahead of the main tunnel top heading face The beginning of the area of small displacements at the far end from the main tunnel face was determined at a 3 mm displacement of a particular measuring point to eliminate the measurement error (the inaccuracy of determining the position of measuring points was less than 2 millimeters [16]). In Fig. 15 the distances from the excavation face to the measuring points at which the measured vertical displacements of a bottom and crown point reached 3 mm and 1 cm, respectively, are plotted along the monitored section of the left tube of the exploratory tunnel. It can be observed that the bottom points moved prior to the crown points. A rather stiff primary lining where the crown point was installed on one hand and no invert at the bottom of the exploratory tunnel on the other can explain such behaviour. Area of recognizable displacements as observed on bottom points reached the length of about 18 - 25 m (2 - 2.5 equivalent tunnel diameters) for a 3 mm displacement and 10 - 20 m (1 - 2 diameters) for a 1 cm displacement in front of the face of the two lane tunnel. Due to a larger excavation cross section in the cavern with simultaneous transition into the fault zone, the uplift of the bottom was registered at larger distances from the top heading face from the chainage km 1.4+95 on. The maximum distance to the excavation face, where the measured vertical displacement of the bottom point exceeded 3 mm, reached 42 m. Similar behaviour was observed for the crown points, only the distances to the face were considerably smaller (10 - 20 m for a 3 mm displacement and 5 - 12 m for a 1 cm displacement in front of the two lane tunnel). 3.3.3 comparison of the displacements in the exploratory tunnel measured during the exploratory tunnel construction with the displacements measured during the main tunnel construction As described above in chapter 3.3.1, the rock mass behaviour in the exploratory tunnel ahead of the face during the main tunnel construction is mainly governed by the orientation of the foliation. Similar dependence can also be observed at the displacements measured during the exploratory tunnel construction. Fig. 16 Chainage [m] bottom p. - 0,3 cm - -*■ - bottom p. -1 cm —•— crown p. - 0,3 cm ■ -•- - crown p. -1 cm Figure 15. Distance of measuring points to the top heading face when the vertical displacement of the bottom and crown point reached 3 mm and 1 cm, respectively, plotted along the left tube of the exploratory tunnel. The scheme of the merging cavern is not in scale. shows different behaviour of the exploratory tunnel due to two different construction activities (construction of the exploratory tunnel and construction of the main tunnel) and consequently two different load cases. When the exploratory tunnel was under construction, the stress change rate due to the excavation was the highest in the area of the excavation face. As the excavation continued, the stress change rate around the same cross section was reduced. Consequently, the displacement rates were high during the initial phase and were close to zero later on. The displacement vector of the crown point tended perpendicularly to the foliation during the initial phase (in the area of large displacements - a blue line marked with letter L in Fig. 16) and parallel to the foliation in the second deformation phase (the area of small displacements - a purple line marked with letter S). Bilinear displacement pattern can only be seen to a limited extent because the absolute values of the measured displacements are small. This phenom- enon became more evident from the displacement measurements ahead of the main tunnel excavation face, since the measured displacements were considerably larger in this case due to larger cross section of the main tunnel compared to the exploration gallery. Moreover in the latter case the entire displacement history was measured. The behaviour ahead of the main tunnel excavation face was just the opposite of the behaviour, observed during the excavation of the exploratory tunnel. The directions of displacement vectors in both cases coincided well when the displacement rates are small (letter S in Fig. 16) and were similar in the area of large displacement rates (letter L in Fig.16). The hypothesis is proposed that the displacements parallel to the foliation dominate when stress change rates are small, while radial displacements govern the behaviour when stress change rates are high. Figure 16. Comparison of the displacements during the execution of the exploration gallery and the displacements measured during the main tunnel construction in the left tube at the chainage km 1.5+17 - cross section together with a face log. 3.3.4 Comparison of the displacements ahead of the face and the displacements within the main tunnel due to the main tunnel construction The comparison of the measured displacements ahead of the face of the main tunnel with the displacements of the main tunnel at the same chainages along the tunnel axis allows the estimation of the portion of the pre-face displacements in total measured displacements and the influence of the orientation of foliation on the orientation of the displacement vectors. The displacement vectors of the exploration gallery and of the main tunnel, caused by the execution of the main tunnel, are plotted in Fig. 17. Similar displacement patterns as described in the previous chapter can be seen. The displacement rates of the monitored cross section were largest some meters ahead (measured on the lining of the exploration gallery) and behind the top heading excavation face (measured on the lining of the main tunnel). Hence, the crown and left sidewall point's displacement vectors tended perpendicularly to the rock mass foliation. The orientation of the displacement vectors changed when the excavation face was far enough from the observed cross section (ahead or behind the face) and the rate of displacements diminished. The magnitudes of the vertical and horizontal displacements of both sidewall points in the exploratory tunnel were approximately the same as the displacements of the sidewall points in the primary lining of the double lane tunnel. The vertical displacement of the crown point was in the exploration gallery than the vertical displacement of the crown point target in the main tunnel and reached about 35% of total measured displacement (marked with a red square in Fig. 18 for the cross section shown). Total measured displacement refers to the sum of displacements measured ahead of and behind the face of the main tunnel. The displacements caused by the exploration gallery execution are neglected. It should be also noted that the targets, where displacements were summed and compared, were not installed at the same places in the observed cross section, as can be seen in Fig. 17. The influence of these simplifications on the assessment of the pre-face portion of total displacements will be studied by numerical analyses. Preliminary numerical study by Jemec showed that the presence of the exploration gallery had limited effect on the behaviour of the main tunnel [17]. Figure 17. Comparison of the displacements ahead of the face and the displacements after the excavation of the main tunnel at chainage km 1.4+44 in the left tube of the Šentvid tunnel - displacement vectors in cross section with face log. Figure 18. The portion of the vertical displacements that occurred ahead of the face for the crown point in the exploratory tunnel and for a point in the main tunnel that is situated above the exploration gallery. In the monitored section of the left tube of the Šentvid tunnel the measured displacements ahead of the face amounted 15% to 45% of the total measured displacements in the same cross section (Fig. 18). A lower portion of the pre-face displacements was observed in stiffer and non-folded rock mass or folded to smaller extent (regions of green and yellow colour in Fig. 18), whilst in more deformable or intensively folded rock mass (regions of red colour) the percentage of the preface displacements was considerably higher. The stated percentage does not take into account the displacement that occurred between the excavation and the first measurement of the newly installed measuring Figure 19. The complete displacement history plot of the crown point in the cross section at the chainage km 1.4+44. section (part 2). The displacement rates in this phase are the highest because of the unsupported rock mass immediately after the excavation and low stiffness of the installed support before the shotcrete hardens. On the basis of the displacement's tangent slope close to the excavation face (some days before and after the excavation of a certain cross section) and time delay between the excavation and the first measurement, we can anticipate the course and the magnitude of the second part of the displacement function's curve (Fig. 19). The estimated portion of the displacements in part 2 is in the range from 10% - 25% of the total displacements. From the presented cases we can deduce that 25% - 70% of displacement occurs before the first measurement of the observed cross section, depending mainly on the stiffness of the surrounding rock mass, construction sequence and the time delay of the first measurement. 4 CONCLUSION Back analyses and interpretation of measured displacements ahead of the excavation face of the main tunnel and perpendicular to its axis and the comparison with the displacements measured during the exploratory tunnel construction as well as the main tunnel construction allow comprehensive interpretation of the rock mass - support response due to the tunnel excavation. A large amount of data was obtained on approximately 250 m long section of the exploratory tunnel of the Šentvid tunnel system. The analyses of these data can provide the knowledge on deformation mechanisms of the rock mass - support system ahead of the tunnel face, which is essential in case of tunnel construction at low overburden under populated area. According to the presented cases one can conclude that the behaviour of the rock mass - support system strongly depends on the foliation orientation if the tunnel is constructed in foliated rock mass of low strength and stiffness like the Šentvid tunnel. The magnitude of the displacements of the exploratory tunnel primary lining due to the main tunnel construction and its portion in all the measured displacements is strongly correlated to the stiffness of the rock mass. The comparison of the displacements ahead of the face with the displacements after the excavation of the main tunnel indicates that 15-45% of the measured displacements occur ahead of the face (less displacement in stiff rock mass that was not folded and more in worse geological - geotechnical conditions). On the basis of the measured displacements in different rock mass types we can assume the strong dependence of the portion of the stress state alternation ahead of the face (parameter Qj) on the stiffness of the rock mass. The geological structure ahead of the face and the size of the tunnel affect the influence area ahead of the face due to the excavation face advance or what is called pre-face domain. The analysis of the measured displacements shows that the small displacement domain can be observed approximately 2 equivalent diameters of the tube ahead of the face and the majority of the displacements occur within a half of an equivalent tube diameter from the face. The value for the parameter xf cannot be deduced from the measured data at present time. Detailed analyses including fitting displacement function to the measured displacements ahead of the face as well as the displacements of the excavated area will define the recommended value for xf. The fitting of the displacement function to the displacements measured ahead of the face allows a judgment on the suitability of the first part of the proposed displacement function and possible modifications of its shape to demonstrate the observed bilinear response of the primary lining in the foliated rock mass. To establish a reliable model for the rock mass behaviour prediction ahead of the face measured displacements will serve as an input to extensive 3D numerical simulations to confirm the suitability of the model. If the proposed model proves to be reliable and accurate, it will be a fast and powerful tool for the prediction of the overall rock mass - support system behaviour. Such tool can contribute to better understanding of the ground behaviour around tunnels with low overburden under populated area [5]. For a given geological structure and known rock mass behaviour on some sections at the same site the displacements of further sections to be excavated can be calculated and therefore sufficient support measures can be designed to comply with the displacement tolerance in the tunnel as well as on the surface. The knowledge on the rock mass behaviour ahead of the face can also serve for the optimization of the support, installed ahead of the face. REFERENCES [1] Rabcewicz L. (1964) The New Austrian Tunnelling Method, Part one, Water Power, November 1964, 453-457, Part two, Water Power, December 1964, 511-515. [2] Vavrovsky, G. M. and Ayaydin N. (1987) Die Bedeutung der vortriebsorientierten Auswertung von Messungen im oberflächennahen Tunnelbau. STUVA-Tagung, Essen. [3] Schubert, W. and Vavrovsky, G.M. (2003) Innovations in Geotechnical on-site Engineering for Tunnels. International Symposium on GeoTechni-cal Measurements and Modelling, Karlsruhe, 35-44. [4] Steindorfer, A. (1998). Short Term Prediction of Rock Mass Behaviour in Tunneling by Advanced Analysis of Displacement Monitoring Data. PhD thesis, Technische Universität Graz. [5] Sellner, P., Grossauer, K., Leitman, R. (2004). How to Predict Surface Movements & Prevent Damages of Surface Structures. Rock engineering - Theory and Practice, Proceedings of the ISRM regional symposium EUROCK 2004 & 53rd Geomechanics Colloquy. Salzburg, 7-9 October, 245-250. [6] Guenot, A., Panet, M., Sulem, J. (1985). A New Aspect in Tunnel Closure Interpretation. Proc. 26th US Symposium on Rock Mechanics, Rapid City, 445-460. [7] Barlow, J.P. (1986). Interpretation of Tunnel Convergence Measurements. MSc thesis, Department of Civil Engineering, The University of Alberta, Edmonton, Alberta. [8] Sellner, P.J. (2000). Prediction of Displacements in Tunneling. PhD thesis, Technische Universität Graz. [9] GeoFit homepage: http://www.geofit.3-g.at/ [10] Klopcic, J. (2004). Visualization and analysis of the displacement monitoring data in tunneling. BSc thesis, University of Ljubljana, Faculty of Civil and Geodetic Engineering. (in Slovene) [11] Schubert, P., Klopcic, J., Stimulak, A., Ajdic, I., Logar, J. (2005). Analysis of Characteristic Deformation Patterns at the Trojane Tunnel in Slovenia. Felsbau, No.5, 25-30. [12] Volkmann, G. and Schubert, W. (2005). The use of horizontal inclinometers for the optimization of the rock mass - support interaction. Underground space use: Analysis of the past and lessons for the future, World tunneling congress, Istanbul, Turkey. [13] Likar, J., Volkmann, G., Button, E. (2004). New Evaluation Methods in Pipe roof Supported Tunnels and its Influence on Design during Construction. Rock engineering - Theory and Practice, Proceed- ings of the ISRM regional symposium EUROCK 2004 & 53rd Geomechanics Colloquy. Salzburg, 7-9 October, 277-282. [14] Žigon, A., Proprentner M., Žibert, M. and Jemec, P. 2006: Šentvid tunnel. Proceedings of the XIII. Danube - European Conference on geotechnical engineering, Ljubljana. 29-31 May 2006, 2nd part, 1025-1030. [15] Čadež, F., Genser, W., Kleberger, J. and Poschl, I. (2004). Šentvid motorway tunnel - Interim results from Slovenia's most recent exploration gallery, Proceedings of the 7th international symposium on tunnel construction and underground structures, Ljubljana, 50-56. [16] Marjetic, A., Ambrožič, T., Bogatin, S., Klopčič, J., Logar, J., Štimulak, A. and Majes, B. (2006). Geodetic measurements in Šentvid tunnel. Geodetski vestnik 50, No.1, Ljubljana, 11-24. (in Slovene) [17] Jemec, P. (2006). Influence of the exploration gallery on the Šentvid two lane tunnel construction. BSc thesis, University of Ljubljana, Faculty of Civil and Geodetic Engineering. (in Slovene) VPLIV POROZNOSTI NA GEOMEHANSKE LASTNOSTI POLŽARICE IZ LJUBLJANSKEGA BARJA_ bojan žlender ln ludvik trauner o avtorjih Bojan Žlender Univerza v Mariboru, Fakulteta za gradbeništvo Smetanova ulica 17, 2000 Maribor, Slovenija E-pošta: bojan.zlender@uni-mb.si Ludvik Trauner Univerza v Mariboru, Fakulteta za gradbeništvo Smetanova ulica 17, 2000 Maribor, Slovenija E-pošta: trauner@uni-mb.si Izvleček Prispevek obravnava mineraloške in fizikalne lastnosti polžarice in njihov vpliv na vrednosti parametrov geomehanskih lastnosti. Polžarica, ime je dobila po vsebovanih fosilih, je tipični sloj Ljubljanskega barja. Je izrazito porozna, zasičena in se nahaja v židkem konsistentnem stanju. V Laboratoriju za mehaniko tal (LMT), Fakultete za gradbeništvo, Univerze v Mariboru je bila izvedena raziskava polžarice s preiskavo mineraloških in fizikalnih lastnosti. Za raziskavo je bil odvzet niz vzorcev na jugozahodni lokaciji Ljubljanskega barja. V Centru za elektronsko mikroskopijo Fakultete za strojništvo, Univerze v Mariboru sta bila raziskana kemična sestava in optični izgled. Kemična sestava je bila določena z visokoločljivim vrstičnim elektronskim mikroskopom SIRION. Optični izgled treh vzorcev polžarice je bil preiskan z okoljskim vrstičnim elektronskim mikroskopom QUANTA 200 3D. Mineralna sestava in zrnatost polžarice sta bila določena na Geološkem zavodu Slovenije. Mineralna sestava je bila določena z rentgenskim difraktometrom. Zrnatost polžarice je bila določena s pomočjo laserskega analizatorja Fritsch particle sizer analysette 22. Specifična površina je bila določena na Kemijskem inštitutu Slovenije, s pettočkovno BET metodo, za meritve je bil uporabljen avtomatski plinski adsorbcijski analizator TriStar 3000, proizvajalca Micromeritics Instrument Corporation, Norcross, ZDA. Ostale fizikalne lastnosti polžarice so bile določene v LMT. Raziskane fizikalne lastnosti kažejo, da je polžarica v naravi zasičena, zelo porozna in skoraj tekoča meljna zemljina. Preiskava stisljivosti kaže, da je polžarica izrazito stisljiva, volumenske spremembe so velike že ob malih spremembah napetosti. Parametri trdnosti imajo značilno nizke vrednosti, ki se ujemajo s predhodnimi raziskavami. Raziskava geomehanskih lastnosti je bila izvedena s poudarkom na sovisnosti fizikalnih lastnosti. Izveden je bil niz triosnih preizkusov polžarice različne gostote oz. vlažnosti in poroznosti. Preizkusi so bili izvedeni v triosnih aparatih na valjastih preizkušancih višine 100 mm in premera 50 mm. Rezultati raziskave kažejo na sovisnost geomehanskih lastnosti in poroznosti. Odnose lahko izrazimo kot funkcije gostote oz. poroznosti ali vlažnosti. Iz rezultatov je razvidno, da so spremembe vodoprepustnosti, konsolidacije in stisljivosti nelinearno odvisne od sprememb poroznosti. Spremembe mehanskih parametrov kot so elastični modul, Poissonov količnik in strižni kot, so pri manjših spremembah poroznosti neizrazite in skoraj linearne. Ključne besede polžarica, triosni preizkus, poroznost, vodoprepustnost, konsolidacija, Youngov modul, Poissonov količnik, strižni kot THE INFLUENCE OF POROSITY ON GEOMECHA-NICRL CHARACTERISTICS OF SNRIL SOIL IN THE LJUBLJANA MRRSH_ bojan zlender and ludvik trauner About the authors Bojan Žlender University of Maribor, Faculty of Civil Engineering Smetanova ulica 17, 2000 Maribor, Slovenia E-mail: bojan.zlender@uni-mb.si Keywords Snail soil, triaxial test, porosity, permeability, consolidation, Young's modulus, Poisson's ratio, friction angle Ludvik Trauner University of Maribor, Faculty of Civil Engineering Smetanova ulica 17, 2000 Maribor, Slovenia E-mail: trauner@uni-mb.si Abstract This article focuses on mineralogical and physical characteristics of snail soil and their influence on parameter values of geomechanical characteristics. Snail soil, which got its name from fossil remains, is a typical layer observed in the Ljubljana marsh. It is distinctly porous, saturated and in a liquid consistency state. Snail soil was investigated for mineralogical and physical characteristics in the Laboratory of Soil Mechanics, Faculty of Civil Engineering of the University of Maribor. Mineral and chemical composition, visual appearance, specific surface and grain property were determined. Physical characteristics show that snail soil is saturated in nature, highly porous and almost liquid. Geomechanical characteristics were investigated for their interdependency on physical characteristics. A series of triaxial tests were performed on snail soil samples of different porosity, density and water content. Cylindrical samples of the height of 100 mm and the diameter of 50 mm were tested using three-axial testing apparatuses. The results of the tests show that interdependency exists between geomechanical characteristics and porosity. These relationships can be expressed as functions of density, porosity or water content. It is evident from the results that changes of the coefficient of permeability, the coefficient of consolidation, and the coefficient of volume compressibility are non-linear with respect to changes of porosity. Changes of mechanical parameters, such as Young's modulus, Poisson's ratio, and friction angle are indistinct and almost linear at lower changes of porosity. 1 INTRODUCTION This article focuses on mineralogical and physical characteristics of snail soil and their influence on parameter values of geomechanical characteristics. Snail soil, which got its name from fossil remains, is a typical layer observed in the Ljubljana marsh. It is distinctly porous, saturated and in a liquid consistency state. The Ljubljana marsh is a wide tectonic sink which was formed two million years ago by a gradual depression of the area. The marsh is located in the south of Ljubljana, at the elevation of 287-290 m above the sea level, and covers the surface of 163 km2. The Ljubljana marsh was inhabited already thousands of years ago. Archeological findings (fascine dwellings) date from the iron, copper, and bronze periods. Later, this marshy area was invaded by the Romans who were the first to start draining the land. The contribution of fascine dwellers from the Ljubljana marsh to a wider cultural space is also proven by the recent archeological finding, i.e. a wheel and the ax of a two-wheel vehicle, which dates from about 3200 A. D. For the present, this archeological finding is known to be the oldest of its type in the world. Today, this area is almost completely drained and urbanized. Yet, the construction of traffic ways and tall buildings presents a great challenge to construction engineers due to the softness of layers below the surface. Geological structure of the marsh is very interesting. Ground water is located immediately below the surface. The surface layer is composed of peat of the thickness of 1 m to 8 m. The dept of the peat is nowadays essentially smaller than in a past, due to the intensive excavations in a first half of the 20th century. Below the peat layer, there is a layer of snail soil of the thickness of a few meters at the borders to more than 10 m in the center of the marsh. The snail soil layer is distinctly porous, saturated with water and of a low-bearing capacity. There are clay and sandy-gravel layers below the snail soil layer. A layer of rocks starts at the depth of some ten meters. Geological structure of the Ljubljana marsh has been studied by numerous experts. The oldest geological documents go back to the middle of the 19th century, when the first geological map was drawn of this region. Later, several other studies were performed. In the second half of the 20th century the researches were focused on geomechanical characteristics of the soil and detailed investigations of physical and mechanical characteristics of snail soil were performed. Strength parameters showing possible static loadings were also determined. To understand geomechanical characteristics of this marshy area it is essential to know rheological characteristics of snail soil. In 1979, in the framework of Trauner's doctoral thesis [1], slow triaxial tests were performed in the Laboratory of Soil Mechanics at the Faculty of Architecture, Geodesy and Civil Engineering of the University of Ljubljana. A similar investigation was later repeated in the Laboratory of Soil Mechanics (LSM), Faculty of Civil Engineering of the University of Maribor. A rheological model of snail soil was set which shows relationships between stresses, deformations and time [2]. The results of investigation were expressed with rheological dependencies, i.e. function relationships between rheological parameters and stress state and deformation, at void ratio e = 2.1. An example of dependencies (Figure 1a) showing the relationship between shear stress t and effective normal stress a, with low value of friction angle q> = 21° and no cohesion c = 0. 10 20 30 40 50 60 70 80 90 100 110120 Effective stress a' [kPa] Figure 1a. Relationship between shear stress t and effective normal stress a'. Fig. 1b shows a typical result obtained for the octahedral strain eg depending on the octahedral shear stress t0 and effective octahedral stress state a' of snail soil. 0,07 V 0,06 Î 0,05 I -§ 0,04 1 u ■a g 0,02 0,03 0,01 0,00 £„ = ( 0 ( 0.96 . This was a relatively short-term phase because of saturation in the nature. The saturated sample was then consolidated at the selected effective isotropic consolidation stress a'c and selected compression stress change Aac. The effective isotropic consolidation stress is expressed as a difference between the cell pressure ac and the back pressure ub. Static loading was performed so that the sample was loaded with the selected compression stress dac or the axial stress Aa = Aa . a z Sixty-two triaxial tests were performed. The investigation was based on a series of tests in which the below conditions varied: - Void ratio e = 2.1 + 1.2 (Ae is calculated), - Initial effective pressures a0'=a'3c = 0, 50, 100, 150 kPa, - Variations of effective pressures Aa' = 50, 100, 150 kPa, - Variations of axial pressures daz (depending on axial deformation). The below pressures were measured during the test: - Cell pressure ac (kPa), - Back pressure ub (kPa), - Pore water pressure uw (kPa), - Axial vertical stress in compression az (kPa), - Axial vertical and radial strain ez (-), er (-), - Volume deformation ev (-). For undrained test ev (-)=0. Constants in the research are: - Specific surface, - Chemical composition, - Mineral composition, - Grain property, - Natural humidity, density and coefficient of porosity, - Liquid limit, plasticity index, consistency index, - Density of a solid mass, - Dry volume weight, - Saturation. The following strength parameters were calculated: coefficients of permeability k (m/s), consolidation cv (m2/s) and volume compressibility mv (kPa-1), constrained modulus Mc (kPa), Young's modulus E (kPa) and Poisson's ratio v (-), or compression modulus K (kPa), and shear modulus G (kPa). Standard oedometer tests and direct shear tests of snail soil were also performed. Snail soil has the void ratio e = 2.1 in nature, which is an extremely high value. Values of mechanical parameters change with lowering the void ratio. Porosity (void ratio e) is in linear correlation with water content w. Density (p) or unit weight (y) is in nonlinear correlation with void ratio e. K = B • (3) where: 1 Yw Dr + (^r • g/100) m w =--e = -w- • e y=—r-—---Y (1) r Ys Y 1 + g ( ) Dr where: yw, ys unit weight of water and solids, Dr relative density of soil, Sr=1 degree of saturation. Figures 4 show the correlation between void ratio e and water content w, and the correlation between void ratio e and density p. Therefore, the snail soil was in liquid or plastic consistency, for different tests. Ks hydraulic conductivity, effective porosity, B, n are constants. The relationship of the coefficient of permeability k vs. void ratio e for all tests is: k(e) = B1 •e" (4) where e void ratio B1, "1 are coefficients. # | plastic | | liquid | 0 wp=37 e=1 p =1,85 < « > > O WlN>0 e=1,6 P =1,65 1 1,2 1,4 1,6 1,8 2 Void ratio c [-] We obtained the expression k = 4 ■ 10-11 ■ e5,5019 for a series of all tests (where e is void ratio). Fig. 5 shows the relationship of the coefficient of permeability vs. void ratio for all tests. Deviation of the results (R = 0.3) from the above function is high because the function does not include stress conditions or pore pressure gradients, respectively. However, a more detailed description of the coefficient of permeability can be made with the functions of void ratio and stress states ag' = 50, 100, 150 kPa or with pertaining pore pressure gradients. | plastic | | liquid | O 9=1,6 p=1,65 > : ❖ : ......................A...................... wp=37 e=1 p=1,85 0 . 0 * T 1 1,2 1,4 1,6 1,8 2 Void ratio e [-] Figure 4. The correlation between void ratio e and water content w, and correlation between void ratio e and density p. 3-1 coefficient of permeability The coefficient of permeability is expressed with equation k = cv • mv • Yw (2) Figure 5. The coefficient of permeability k vs. void ratio e. The coefficient of permeability of snail soil in nature is k = 2 •10-9 m/s. The value of the coefficient of permeability decreases with decreasing void ratio e as shown in Fig. 5. The permeability can be expressed similar to a known expression of Ahuja et al. [3]. Figure 6. The coefficient of permeability k vs. water content w. If the relationship of the coefficient of permeability k vs. void ratio e (or density p, or water content w) is expressed in the logarithmic form, we can see that it is almost linear. Figure 6 shows this relationship as the coefficient of permeability k vs. water content w. 3.2 coefficient of consolidation The coefficient of consolidation cv is expressed using equation from BS 1377 [4] c _ 1,65- D2 A - f,„„ (5) where: 1.E-03 , 8.E-04 •I 3, 6.E-04 S * ^ S * è 4.E-04 I 1 fe I 2.E-04 'S I 8 O.&OO 0 0 ..........o° «..... 0 ' ^ % ^^ : ° ♦ 0 ................!.............. y = 9E-05X3-5005 ; 1,200 1,400 1,600 Void ratio e [-] 1,800 2,000 Figure 8. The coefficient of volume compressibility mv vs. void ratio e. D diameter of the specimen (mm), X coefficient which depends on the drainage, tigg time of primary consolidation. The relationship of the coefficient of consolidation cv vs. void ratio e could be expressed similarly to the coefficient of permeability k. Fig. 7 shows the relationship of coefficient of consolidation c vs. void ratio e for all v c tests. We obtained the expression cv = 3 -10-8- e1,464 (e is void ratio). 1,&07 1, 8.E-08 it I 6.E-08 I S 4.E-08 B s te 2,E-08 0.&00 ❖ -O—O- 0 ❖ 0 0 t ^ -— »0 > 0 0 ........♦............... 0 > y = 3E-08X1'4644 1,600 Void ratio e [-] 1,800 Figure 7. The coefficient of consolidation cv vs. void ratio e. 3.3 coefficient of volume compressibility The coefficient of volume compressibility mv is expressed with equation Ae Ae„ (1 + e)-Act ' Act ' (6) Fig. 8 shows the relationship of the coefficient of volume compressibility mv vs. void ratio e for all tests. We obtained the expression mv = 9 ■ 10-5 ■ e3,5005 . The logarithmic form of relationship of volume compressibility mv vs. water content w shows linear correlation log. mv = 3 ■ 10-10 ■ w3,5005. 1.E-03 Water content w [%] Figure 9. The coefficient of volume compressibility my vs. water content w. 3.4 strength parameters To determine the relationship of strength parameters vs. void ratio an insufficient number of tests were performed therefore the results are unreliable. Fig. 8 shows Young's modulus E vs. void ratio ec. We can see that strength does not substantially increase with increasing density or decreasing porosity; a greater difference can be only seen at higher changes of density or porosity. Poisson's ratio is v = 0.4 at void ratio e = 2.0 and decreases for 0.03 at void ratio e = 1.4. 1,6 i,i Void ratio e [-] Figure 10. Young's modulus E vs. void ratio e. If the relationship of the Young's modulus E vs. water content w is expressed in the logarithmic form, we can see that it is almost linear (Fig. 10). mv _ •S 1000 - o S .ô- X y = 7B08X"3-1' Water content w [%] ACKNOWLEDGMENTS This research is being financed by the Slovene Ministry of Higher Education, Science and Technology. The financial support of DARS, the Slovene Motorway Company, and the company Prevent is also gratefully acknowledged. Figure 11. Young's modulus E vs. water content w. REFERENCES The same is true for shear properties. Fig. 12 shows the relationship of friction angle q> vs. unit weight y. We can see that it increases almost linearly at lower changes of density or void ratio, respectively. 31 29 E 27 25 60 g 23 I 21 I 19 17 15 15 15,5 16 16,5 Unit weight y [%] O y = O.OOx3-79 17,5 [1] Trauner, L. (1982). Applicability of theory of elasticity for foundation design, Doctoral thesis, University of Ljubljana, 479 p., Ljubljana [2] Trauner, L. et al. (1982). Structure soil interaction, Research report, University of Maribor, 67 p., Maribor [3] Ahuja, L. R., et al. (1989). Evaluation of spatial distribution of hydraulic conductivity, using effective porosity data, Soil Science 148, p. 404-411 [4] British Standards Institution (1999). Consolidation, BS 1377, Part 8, London Figure 12. The friction angle f vs. unit weight y. 4 CONCLUSIONS This article focuses on the investigation of snail soil and the research of its mineralogical and physical characteristics. Geomechanical characteristics were investigated for their dependence on physical characteristics. A series of triaxial tests of snail soil of different density, porosity and water content was performed. The results of the tests show that geomechanical characteristics depend on porosity. The relationships were expressed as functions of porosity. It is evident from the results that changes of the coefficient of permeability, the coefficient of consolidation, and the coefficient of volume compressibility are nonlinear with respect to void ratio. Changes of mechanical parameters such as Young's modulus, Poisson's ratio and friction angle are indistinct and almost linear at lower changes of porosity. METODE ZA KONTROLO PRONICANJA VODE SKOZI PREGRADE IZ VALJANEGA BETONA Z MERITVAMI VODOTESNOSTI IN DRENAŽ yueming zhu, stephan semprich, erich bauer, cuiping yuan ln dongmei sun o avtorjih Yueming Zhu Institute of Hydrostructure Engineering, Hohai University Nanjing, Kitajska E-pošta: ym-zhu@163.com Cuiping Yuan Hohai University Nanjing, Kitajska Stephan Semprich Institute of Soil Mechanics and Foundation Engineering, Graz University of Technology Gradec, Avstrija E-pošta: stephan.semprich@tugraz. Dongmei Sun Hohai University Nanjing, Kitajska E-pošta: sundongmei@126.com Erich Bauer Institute of Applied Mechanics, Graz University of Technology Gradec, Avstrija E-posta: erich.bauer@tugraz.at Izvleček Kitajska je v zadnjih nekaj letih močno razvila svojo tehnologijo izgradnje jezov iz valjanega betona (beton RCC). Trenutno se ta tehnologija uspešno uporablja pri izgradnji izredno visokih gravitacijskih jezov, pa tudi pri izgradnji srednje visokih in visokih ločnih jezov. Na Kitajskem trenutno projektirajo ali gradijo kakšnih sto jezov iz valjanega betona. Eden glavnih tehničnih problemov pri takšni gradnji je razumevanje pronicanja vode skozi valjani beton ter potreba po razvoju ustrezne teorije in metod za nadzor pronicanja. Prvi avtor tega članka zato že več kot deset let sodeluje v obsežni študiji o lastnostih valjanega betona in metodah za nadzor pronicanja. Študija podaja teoretične osnove ter rezultate terenskih raziskav o pronicanju vode, ukrepih za doseganje vodotesnosti, drenaži, optimalnem načrtovanju in nadzoru pronicanja. Rezultate te študije smo koristno uporabili pri izgradnji in pri povratnih analizah številnih jezov. Članek podrobno opisuje raziskovalne dosežke študije, teoretske osnove ter njihovo praktično uporabo pri visokih gravitacijskih jezovih iz valjanega betona. Ključne besede gravitacijski jez, valjani beton, RCC, pronicanje, anizotropna prepustnost, drenaža METHODS FOR CONTROL OF SEEPAGE IN RCC DAMS WITH WATERTIGHT AND DRAINAGE MEASURES yueming zhu, stephan semprich, erich bauer, cuiping yuan and dongmei sun About the authors Yueming Zhu Institute of Hydrostructure Engineering, Hohai University Nanjing, P.R. China E-mail: ym-zhu@163.com Stephan Semprich Institute of Soil Mechanics and Foundation Engineering, Graz University of Technology Graz, Austria E-mail: stephan.semprich@tugraz.at Erich Bauer Institute of Applied Mechanics, Graz University of Technology Graz, Austria E-mail: erich.bauer@tugraz.at Cuiping Yuan Hohai University Nanjing, P.R. China Dongmei Sun Hohai University Nanjing, P.R. China E-mail: sundongmei@126.com Abstract The technologies for construction of roller-compacted concrete (RCC) dams have been considerably developed during recent years in China. At the time being, they have been successfully applied to the constructions of even extreme-high gravity dams and medium to high arch dams. There are a few of hundreds of RCC dams (RCCD) under design and/or construction in China. One of the main concerned technical problems according to the construction is about the understanding of the property of seepage in RCCDs and the relevant theory and methods for the control of the seepage. In order to overcome the problem, the senior author has been engaged in a wide study on the property and methods for control of seepage in RCCDs for more than 10 years. The property of seepage, measures for watertightness and drainage, optimal design and construction schemes for control of seepage in the dams have been essentially understood either in theory and practice. The results have been applied for the construction and the backanalysis of several dams. The paper describes the research findings in detail with respect to the theoretical fundament and their application for a high RCC gravity dam. Keywords gravity dam, Roller-compacted concrete, RCC, seepage, anisotropic permeability, drainage 1 INTRODUCTION The Horizontal Layer Method used in roller-compacted concrete (RCC) construction has been replaced by the Slope Layer Method, which importantly reduces the area of exposed young RCC and increases RCC placing rates. The intrinsic permeability of RCC is very low (it equals 1.0x10-9 to 1.0x10-12 cm/s), whilst the permeability of the lift surface including the joint lift surface is proportional to the cubic width of a hydraulic joint. Theory and engineering practice have proved that the lift surface is the main pathway of seepage, in which leakage generally occurs. According to the present technology for constructing of RCC dams, the body of a RCC dam is considered to be a strong anisotropic medium with more than 2 to 4 or even 6 to 7 orders of magnitude (for example Willew Creek RCC gravity dam, USA) of ratio of the tangential major coefficient of permeability to the normal major coefficient of permeability of the layer and lift joint surfaces. This causes the seepage properties of RCC dams to be totally different from those of conventional concrete dams. Since the early development stages of RCC dam construction, there have been a lot of problems with seepage due to the lack of knowledge about the property of seepage in RCC dams. One problem, for example is an extremely high position of the exit point on the downstream face, so that exit-flowing discharge may be higher than expected. This is caused by the anisotropic hydraulic resistance according to the dam structure constructed by layers. So, a lot of dams were required to increase their waterproof capability by grouting during the beginning period of operation. Such was the case with Willew Creek RCC dam and Chinese Shimantan RCC gravity dam of the height of 40.5 m. This paper analyses in detail optimal design schemes for the control of seepage. The analysis is based on a wide and in-depth study on the property and methods for the control of seepage in RCCD and on the senior author's experiences in studying characteristics of permeability and control techniques to handle problems of seepage for RCC dams. 2 PROPERTIES OF SEEPAGE OF RCC AND RCCD The system of RCC or RCCD consists of body layer of concrete with homogeneous and isotropy property, a layer and a lift interface of anisotropy property which becomes an inhomogeneous multi-laminated medium. Actual properties of seepage of RCC and RCCD are influenced by the properties of the body of concrete, which can be described by Darcy's coefficient of permeability kRCC , and the properties of the layer and the lift interface which is related to the width of a hydraulic joint, roughness of the joint, connectivity of the joint, stress and strain behaviour of the layer surface, load history, etc. The engineering practice has already proved that the layer and the lift interface are the main pathway of seepage and represent weak surfaces according to tension and sliding. In general, RCC and RCCD can be considered as strongly anisotropic medium. An experiment recorded in literature [1, 2] presents the model of an in-situ concrete block connected in parallel and in series within the Longtan RCC project, which is shown in Fig. 1. The experiment shows that the ratio of average coefficient of permeability in the parallel model to that in the series model is about 1 to 4 orders of magnitude. And the seepage flow mainly exits through the layer and the surrounding region of great skeletal material, which proves that the layer and the lift interface are main pathways of seepage [3]. Figure 1. Parallel and series models of a rectangular RCC block in the permeability test. Gong et al. [3] show the results of in-situ water pressure test, which includes the tests on Willew Creek RCC dam in the USA, Chinese Guanyinge and Tongzijie projects. The tangential major coefficient of permeability of the layer in Willew Creek RCC dam in the USA even surpasses 3x10-3 cm/s, and the average value of the tangential major coefficient of the layer permeability is in the range of 1x10-3 to 1x10-8 cm/s. So, heavy leakage occurs through layers and the location of the exit-flowing line is very high. Compared to the tangential major coefficient of layer permeability, the normal major coefficient of layer permeability is the same as the coefficient of permeability of the body of RCC. The properties of seepage of RCC and RCCD mainly depend on the tangential property of seepage. The system of RCC and RCCD should be considered as strongly anisotropic, and the ratio of anisotropy may reach 1 to 4 or even more orders of magnitude. Theoretically, the properties of flow through layer and lift joint surfaces agree with the properties of fissure flow. The properties of seepage flow through layer and lift joint surfaces in RCC and RCCD can be usually described by the cubic law of an advective fissure flow, which has been proved by home and foreign engineering experts and which is in accordance test with results in various countries [3, 4]. The property of RCC considered as inhomogeneous multi-laminated medium can be described by tangential and normal major coefficients of the permeability of the layer which depend on the body coefficient of permeability, cubic law and the thickness of the layer. Eqs. (1) and (2) show the tangential and normal major coefficients of the permeability of the layer. kt = i[(B - df )kRcc 3 ] 12^ B k = f ' RCC Bk B - d (2) f where kt and kn are average tangential and normal major coefficients of permeability of the layer, respectively, B denotes the thickness of the layer including the body thickness of the layer and the width of the hydraulic joint df , kRCC denotes body coefficient of permeability of RCC, f denotes the kinematic viscosity of water with f » 0.013 cm2/s at 10° C. Because df is much smaller than B , the normal major coefficient of the permeability of the layer mainly depends on the body coefficient of permeability of RCC and also on the coefficient of bedding cushion, when there is one, whilst the tangential major coefficient of permeability mainly depends on the width of a hydraulic joint. Figure 2. Cross section of seepage control measures for RCCD. Due to a strong anisotropic property of RCC, the seepage control measures for RCCD become more difficult than conventional concrete. Fig. 2 shows the general layout of measures for watertightness and material zones. It will be a reliable anti seepage structure of a distorted concrete layer with a thickness of 0.3 to 1.0 m on the upstream face, which plays an important role of water-resistance in the front of the dam. Other forms of anti seepage structure on the upstream face of RCCDs are available, such as a RCC layer, a reinforced concrete slab, an asphalt concrete layer, and a combination of different anti-seepage structures. The majority of all projects under construction or under design adopt a distorted concrete layer as the anti seepage structure. Some infiltrating flow through the anti-seepage structure will be drained by drainage holes behind a distorted concrete layer, in which the uplift pressures in the layer and lift surface are almost zero. In order to prevent the seepage flow from exiting on the downstream face, a thin distorted concrete layer is also required on the downstream face. When the upstream water level is very high, a curtain of drainage holes hanging upward the datum plane is needed. Sometimes horizontal drainage holes are alternatively arranged in the layer and the lift surface additionally helps drainage and pressure relief. The principles of seepage control for RCCDs foundation mostly agree with conventional concrete dams which mainly depend on a curtain of watertightness and drainage holes. 3 NUMERICAL ANALYSIS ON THE RCCD SEEPAGE FIELD WITH FEM 3.1 numerical model of seepage field Due to special inhomogeneous and strongly anisotropic multi-laminated structures in RCCDs, the system of RCC should be considered as fractured rock masses medium according to the property of the layer seepage, which can be simulated by an equivalent continuum model, a discrete fractured network model and a dual porosity model [4-7]. Equivalent continuum models based on the balance equation of mass consider layer and lift surfaces as equivalent continuum porous media, which are analyzed by a mature theory of continuum porous media. But the equivalent seepage discharge and seepage pressure cannot be obtained with this method. A discrete fractured network model assumes that the body of RCC is impervious, so that the properties of RCC mainly depend on the widths of hydraulic joints, location and connectivity of the layer, and the coefficient tensor of permeability is determined by the cubic law and Darcy's law. This method is able to actually simulate the behaviour of seepage flow between layers, but due to the uncertainty of the joint and lift surface, it is more difficult to simulate fractures. The objects within a dual porosity model involve seepage behaviour both in porous media and fractured media. Combining the merits and faults of three models mentioned above, the authors proposed inhomogeneous multi-laminated media elements model, which unifies the equivalent continuum model and the non-continuum mixed model, and which has the advantages of an equivalent continuum model involving fewer elements needed in the calculation [8]. 3.2 inhomogeneous multi-laminated media element An inhomogeneous multi-laminated media element model assumes that the dam body consists of a body layer of roller-compacted concrete, a lift interface, a lift joint interface of the course of one continuous construction of lifts, a bedding plane and normal concrete materials, which become inhomogeneous multi-laminated media elements. The element conductivity matrix is expressed with Eq. (3). kj = /-'1 /-'1 /-'1 BTKBJd^vdC = E //" (£ /- BfKtBjU\d^dv)d( (3) Here we introduce the integral variable transformation shown in Eq. (4). Z_ Zk Ck-1 z' + Zk Ck-1 (4) Because the normal major coefficient of the permeability of a RCC body and the width of the joint are so small that normal head loss in the joint lift surfaces is almost zero, the seepage flow through joint lift surfaces can be considered as a two dimensional flow model (index f). We then have Eq. (6). d dxf kf dL k dxf =0 i, j = 1,2 (6) where xf is the corresponding local coordinate, kf is a two dimensional coefficient tensor of permeability in the planar element without the thickness model, and kfes is the element conductivity matrix which can be calculated from the below Eq. (7): f 1 81VL+2kf N N + f N N ds dxf dxf dxf dxf dxf dxf (r, s = 1-- - m) (7) where sf is a subdomain of a joint lift surface element, Nr and Ns are interpolating functions of the joint lift surface element, m is the number of nodal points in the joint lift surface element. According to balance equation of mass [11], we obtain the below FEM governing equation for a seepage field So we have Eq. (5). ^ (QRCC +Qf ) = 0 i = 1,2,- ■,n (8) ktJ _ J- Zk Zk-1 (J-J- BfKkB] Jdd'jd( (5) where n is the total number of nodal points, Q, k_1 f^i o-ro r»r»/'1al flnvoc nAint 1 o-ro /~rv Here N is the number of layers of inhomogeneous multi-laminated media, Kk is the coefficient tensor of permeability of the k th layer, Eqs. (9), (10) and (11), B{, Bj and |/| are functions of the local coordinates of £ , n and Z . 3.3 planar element without thickness for fissures Bearing in mind numerous lift interfaces and lift joint interfaces and occurrence of horizontal cracks and cleavage cracks in case of emergency of a dam, the authors proposed the planar element without the thickness model ([9] and [10]), which can be used to simulate the exact behaviour of the seepage flow in RCC and which has been proved to be a successful model up to now. and Qf are nodal fluxes of the point i which are contributed by a 3-D equivalent continuum media element of RCC and a 2-D planar element without the thickness element. 3.4 coefficient tensor of permeability in the rccd seepage field In the study of seepage properties in RCCDs, if the layer of RCC is horizontal, the coefficients of permeability can be described by major coefficients of permeability shown in Eqs. (1) and (2). In a sloped layer the direction of major coefficients of permeability are not the same as a coordinate axis and the properties of seepage should be expressed by the coefficient tensor of permeability shown as Eqs. (9), (10) and (11). Eq. (9) is suitable for the layer incline to the upstream, whilst Eqs. (10) and (11) are suitable for the layers' incline to the left bank and the right bank, respectively. [ K ] = k r cos2 Q + sin2 Q 0 0 r 1 ( r -1) sin2Q (9) (r -1)sin2Q 0 rsin2 Q + cos2 Q r 0 [K] = kn 0 r cos2 Q + sin2 Q 1 1 (1 - r ) sin2Q 0 -( 1- r ) sin2Q r sin2 Q + cos2 Q 2 (10) upstream face, and seepage control treatment is done on the lift surface within the width of 2 m (i.e. to spread a layer of mortar of the thickness of 1.0 ~ 1.5 cm in time). Different projects will take different schemes of control of seepage. Temperature control and crack prevention are very important on the upstream face to prevent the occurrence of penetrating cracks. Quality of structure of seepage control is more important than its width. Uplift pressure tolerances in the lift surface, the joint lift surface and the datum plane are shown in Figs. 3 and 4, where Hu and Hd are depths of upstream and downstream respectively and y unit weight of water. [ K ] = k 0 r cos2 Q + sin2 Q -(r-1)sin2Q (11) 0 1 (r -1) sin2Q r sin2 Q + cos2 Q where 9 is the slope angle, kt and kn are tangential and normal major coefficients of permeability, respectively, r = is the ratio of tangential major coefficient of permeability to normal major coefficient of permeability, [ K ] is the coefficient tensor of permeability. 4 THE DESIGN METHODS FOR WRTERTIGHT RND DRRINRGE MERSURES 4.1 watertight structures on the upstream face Lift surfaces on the upstream face are probably a pathway of leakage, so it is important to take watertight measures in the upstream. The distorted RCC and 2-grade RCC are mostly used as watertight measures on the upstream face in present RCCDs. 2-grade RCC represents a defined RCC quality, which is higher than the 3-grade RCC. Based on the authors' experiences about methods of seepage control in RCCDs, the following schemes for the control of seepage are proposed. First, a thin layer of distorted RCC and additional 2-grade RCC are arranged on the upstream face, and seepage control treatment is done on the lift surfaces within the width of 2m. Second, a thin layer of distorted RCC and additional 3-grade RCC are arranged on the Figure 3. Cross section of design uplift pressure in the dam body. Figure 4. Cross section of design uplift pressure on the datum plane. 4.2 watertight structures on the downstream face Due to the strong permeability of the lift surface on the downstream face of RCCDs, the location of the exit-flowing line may be very high and the leakage of seepage may be so great that it will endanger stability safety of the dam. A layer of distorted RCC with a thickness of 0.30~0.50 m will be arranged on the downstream face. Inside the distorted RCC layer, the seepage control treatment will be performed to the depth of 1 to 2 m of the r lift surfaces below the water level of downstream; alternatively; 2-grade RCC can follow inside the distorted RCC layer. 4.3 drainage holes curtains in the dam body Drainage holes are required both in a conventional dam and a RCC dam to discharge the seepage flow. The curtains of exit-flowing drainage holes will be arranged in the zone of 3-grade RCC following 2-grade RCC in the upstream region, in which the intervals of drainage holes are about 4 to 5 m. In the downstream region of a 2-grade RCC zone or within the RCC, treated by seepage control, a curtain of exit-flowing drainage holes can be also arranged, in which the interval between drainage holes is recommended to be 6 m. For a high dam with a wide datum plane, the curtains of overflowing drainage holes are needed on the datum plane near the upstream region. 4.4 bedding plane of concrete on the datum plane Due to cracks generated frequently, difficulties in construction and more cement used in conventional concrete, the bedding plane built of conventional concrete has been replaced by bedding plane built with 2-grade RCC of the thickness of 0.5 m to play the role of providing flat and resisting water. The bedding plane and watertight structures together form a watertight structure. 4.5 watertight structures and drainage measures on the datum plane Watertight structures and drainage measures are also required in the zones of the dam heel and the dam toe on the datum plane. Engineering practices have proved that curtain's relative watertightness just needs to be 3 times lower than of the surrounding rock masses. The grout curtain must be complete and the curtain drainage holes should be unblocked. The interval of drainage holes on the datum plane is 4 to 5 m, the same as in the dam body. 4.6 transverse joint seal and watertight structures on sides of joints Compared to conventional concrete dams, the RCCDs need better transverse joint seal because there are lift surfaces in RCCDs. In order to prevent water from infiltrating into the dam through the sides of the transverse joint, watertight structures on the sides of the joint are strongly required. This can be realized by a distorted concrete layer of a thickness between 0.3 and 0.5 m and should be connected very well with the bedding plane. 4.7 design principles for the watertight structure and drainage measure Briefly, the design principle of a watertight structure and the drainage measure is to resist in front and to drain at the back in the upstream region of the dam as well as in the zone of the dam toe in dam foundation, whilst in the downstream region of the dam the principle is to drain in front and to block at the back as well as in the zone of the dam heel in dam foundation. Here, to resist in front in the upstream region means to arrange a layer of distorted RCC and 2-grade RCC or the zone of lift surface and lift joint surface treated by seepage control method of the width of 2 to 3 m; and to drain at the back means to arrange the main curtain of drainage holes. To drain in front in the downstream region means to arrange the main curtain of drainage holes near the upstream, and to drain at the back means to arrange a layer of distorted RCC. In the zone of the dam toe in dam foundation, to resist in front means to arrange the main grouting and the watertight curtain, and to drain at the back means to arrange the main curtain drainage holes behind the grouting and watertight curtain. In the zone of the dam heel in dam foundation, to drain in front means to arrange the curtain of drainage holes behind the watertight curtain near downstream, and to resist in the back means to arrange the watertight curtain near downstream. In fact, the function of water-resisting of the watertight curtain and the function of drainage and pressure relief of the curtain of drainage holes are not independent. If there is only a watertight curtain and no curtain of drainage holes in dam foundation, the uplift pressure will be very high unless the permeability of the watertight curtain is very small, for example 2 orders of magnitude smaller than that of surrounding rock masses. This is difficult to perform in engineering practices. If there is only a curtain of drainage holes and no watertight curtain in dam foundation, though the uplift pressure can be greatly reduced, the rock masses in foundation may be destroyed by a great hydraulic gradient surrounding the drainage holes, which is absolutely irrational for the long-term safety of dam. 5 APPLICATION 5.1 study on the entire seepage field We have taken the 6th segment of Guangzhao RCCD as a case study, which is located in the middle reaches of the Longtan River in the Guizhou province. On the face of upstream, there is a layer of distorted concrete of different thickness along the height, 0.8m above the elevation of 615.0 m and 1.0 m below the elevation of 615.0 m. The watertight structure in the upstream region is composed of a layer of distorted concrete and additional 2-grade RCC and 3-grade RCC, except for the zone above the elevation of 710.0 m, where there is a layer of distorted concrete and an additional 3-grade RCC. The depth of 2-grade RCC is in the range of 3 to 13m according to the water head on the face of upstream. There is a drainage curtain in the dam body, and the watertight curtain and the main drainage curtain are arranged in the zone of the dam toe because the water level in downstream is lower than the elevation of the datum plane. The normal water levels are 745.0 m and 583.5 m in the upstream and downstream, respectively. Empirical coefficients of permeability of various watertight materials in the dam, the watertight curtain in dam foundation and the grout curtain are shown in Table 1. According to the seepage control methods proposed in this paper, Fig. 5 and Fig. 6 show the distributions of water head contour lines and uplift pressure head contour lines, respectively, in a short form. The entire seepage field presents exact seepage property and obvious regularity, which indicates that the seepage control method proposed in this paper plays the role of Figure 5. Water head contour lines. Figure 6. Uplift pressure head contour lines. Table 1. Coefficients of permeability of body material and various watertight materials. Type of concrete Body coefficient of permeability (cm/s) Normal coefficient of permeability of lift surface (cm/s) Tangential coefficient of permeability of lift surface (cm/s) notes Conventional concrete 1x10-10 anisotropic Distorted concrete 1x10-9 anisotropic 2 grade RCC 1x10-9 1x10-9 1x10-7 3 grade RCC 1x10-9 1x10-9 5x10-7 Watertight curtain 1x10-5 isotropic Consolidated grouting 1x10-4 isotropic resisting in front and draining at the back. In the dam body, the potential of high water head is mostly lost through the isotropic layer of distorted concrete. Though the coefficient of permeability of distorted concrete is the same as the normal major coefficient of permeability of 2-grade RCC and 3-grade RCC behind the layer of distorted concrete, it is watertight compared to the tangential major coefficient of permeability of 2-grade RCC, because the coefficient of permeability of layer of distorted concrete is 2 orders of magnitude smaller than the tangential major coefficient of permeability of 2-grade RCC. Contour lines in the layer of distorted concrete are probably parallel to the face of upstream. In the layer of distorted concrete above the datum plane about 1 m the potential of water head is reduced about over tens meters, in which the cracks always occur and quality of the layer of distorted concrete should be paid more attention. In the adjacent zone of 2-grade RCC and 3-grade RCC the contour lines have a distinctly horizontal kink towards downstream, which is caused by the fact that hydraulic gradients in 2-grade RCC and 3-grade RCC are much smaller than those in the layer of distorted concrete. Due to the easiness of seepage flow through 3-grade RCC, the quality of a watertight structure ahead 3-grade RCC should be paid more attention to ensure the safety of RCCDs. Some seepage flow moves around the drainage holes, and it will exit from the face of downstream if there is no watertight structure arranged. Fig. 5 shows a layer of distorted concrete on the face of downstream shown in enlargement. Due to no resistance on this seepage flow in the lift surface and lift joint surface, the value of uplift pressure will be near zero. A drainage hole is another useful seepage control structure, by which almost all of the seepage flow is discharged. So, in order to ensure that the drainage hole performs successfully, the quality of the drainage hole should be paid more attention to the construction phase and the drainage hole should be checked regularly during the operation phase. The uplift pressure contour lines at the elevation of 685.0 m, 655.0 m and 625.0 m are shown in Fig. 6. In these typical elevations, the uplift pressure is low, and the drainage curtain in the dam body, the grout curtain and the main drainage curtain in dam foundation play an important role for pressure relief. Briefly, the seepage control methods proposed in this paper work very well. Different watertight structures play different roles and also depend on each other. From the results mentioned above we can see that the layer of distorted concrete and drainage holes play an important role of water-resisting and pressure relief. A big deal of discharge will be drained by the drainage holes. The uplift pressure will be very high if there are no drainage holes. From the distribution of uplift pressure on the datum plane, we can conclude that the grout curtain and the drainage curtain in dam foundation are very important for water-resisting. 5.2 sensitivity analysis on seepage control methods 5.2.1 water-resisting Function of 2-grade rcc Fig. 7 shows the computed result when the 2-grade RCC on the face of upstream is removed on the basis of seepage control structures mentioned above, from which we can conclude that there are no obvious changes in the entire seepage field. We can so come to the same conclusions as those obtained by engineering practice. First, the water-resisting function of 2-grade RCC can be ignored. Second, the layer of distorted concrete on the face of upstream only works in the function of water resisting. Third, the arrangement of 2-grade RCC will make the construction more difficult, reduce the construction rate and make the advantages of RCCD fade away. Fig. 8 shows the distributions of uplift pressure heads, from which the same conclusions can be obtained. Figure 7. The contribution of hydraulic head contour lines in the dam body when 2-grade RCC is removed. so the uplift pressure contour line is horizontal in the layer of distorted concrete. Figure 8. The distribution of uplift pressure head contour lines in different elevations when 2-grade RCC is removed. 5.2.2 The Study of Cracks in the Dam Body Cracks in the dam body are of the key importance for the safety in hydraulic engineering. Figs. 9 and 10 show the computed results when there is a vertical cleavage crack in the layer of distorted concrete on the face of upstream. The width of this cleavage crack is 0.2 mm and it is located in the middle plane between two drainage holes. Figs. 11 and 12 show the computed results when there is a horizontal penetrating crack at intervals of 20 m only in the layer of distorted concrete on the face of upstream. The width of this horizontal penetrating crack is 0.1 mm. When there is a vertical cleavage crack in the layer of distorted concrete on the face of upstream, the function of water-resisting of the layer of distorted concrete disappears and the main watertight structures are 2-grade RCC and drainage holes behind the layer of distorted concrete. The hydraulic head contour lines in 2-grade RCC are parallel to the face of upstream, and the uplift pressure head in the dam body is still small. But due to the anisotropic property of 2-grade RCC, the uplift pressure in 3-grade RCC is a somewhat higher compared to the distribution of uplift pressure shown in Fig. 6. Fig. 11 shows the distribution of hydraulic head contour lines when there are several horizontal cracks on the face of upstream. The values of uplift pressure meet the demands desired. Except for the zones near the cracks, the influence of the cracks on the seepage field can be ignored. Under the watertight function of 2-grade RCC and drainage holes, the uplift pressure head contour lines are horizontal and small enough. Because there is a horizontal crack in the elevation of 685.0 m, the watertight function of the layer of distorted concrete vanishes, Figure 9. Distribution of hydraulic head contour lines when there is a vertical cleavage crack on the face of upstream. Figure 10. Distribution of uplift pressure head contour lines in different elevations when there is a vertical cleavage crack on the face of upstream. Fig. 9 to 12 show the computed results which take into account of various adverse conditions. We can conclude that under the protection of other watertight structures there are no fatal uplift pressures in the seepage field, and the dam is in a temporary safe state. But we are all aware that the tendency of cleavage cracks and horizontal cracks is difficult to predict, so we have to pay more attention to the temperature control and cracks prevention measures, especially to the quality of the layer of distorted concrete. Figure 11. Distribution of hydraulic head contour lines when there is a horizontal penetrating crack at intervals of 20m in the layer of distorted concrete on the face of upstream. Figure 12. Distribution of uplift pressure head contour lines in different elevations when there is a horizontal penetrating crack at intervals of 20m in the layer of distorted concrete on the face of upstream. 6 CONCLUSIONS AND SUGGESTIONS This paper generalizes the properties of seepage in RCC and RCCDs. Constructive suggestions of seepage control methods in various zones of the dam are proposed, and they have been applied to engineering practice. Because the RCC can be considered as inhomogeneous multi-laminated medium with strong anisotropy, the property of seepage of RCC are mainly depends on the property of the lift surface, in which the coefficient of permeability is 2 to 3 orders of magnitude higher than that of the RCC body. In construction stages the quality of the lift surface should be paid more attention. Due to the anisotropy of RCC, the property of seepage of RCC is more difficult than that of conventional concrete. The design principles of watertight structures in RCCDs are to resist in front and drain at the back. On the face of upstream the arrangement of a watertight structure is strongly required, such as reinforced concrete slab, a conventional concrete slab and other forms. In present engineering practice the layer of distorted concrete and additional 2-grade RCC are recommended: the layer of distorted concrete plays the primary role in the watertight function, whilst 2-grade RCC plays the secondary role in the watertight function. In order to drain a small amount of seepage flow through watertight structures, drainage holes should be arranged behind the watertight structure in the upstream. In order to ensure that drainage hole play a permanent role in pressure relief, the diameter of a drainage hole should be greater and the distance between two drainage holes can reach 4 to 6m. In order to prevent the seepage flow from exiting too high on the face of downstream, a layer of distorted concrete is strongly required. The design of watertight control in the dam foundation is the same as that in conventional dam foundation, in which the grout curtain and main drainage curtain are absolutely needed. Based on the study of RCC for over ten years, the design of seepage control methods and the numerical model in RCCDs have been completely solved. The problem of temperature control and crack prevention is still to be researched in future. references [1] Su, B. et al. (2002). Statistical features of permeability coefficients of Jiangya RCCD core samples. Journal of Hohai University, Vol. 30(2), 1-5. [2] Su, B. et al. (2000). Research of permeability of roller compacted concrete of Jiangya Dam. Journal of Hohai University, Vol. 28(2), 7-11. [3] Gong, D., Zhu, Y. (2002) Research on seepage of RCC Dams. International Journal Hydroelectric Energy, Vol. 20(2), 23-25. [4] Wittke, W. (1990). Rock Mechanics - Theory and Applications with Case Histories. Springer-Verlag Berlin Heidelberg. [5] Schrader, E.K., Namikas, D. (1988). Performance of roller compacted concrete dams. Proceedings of the Sixteenth Congress on Large Dams, Vol. 2, 339-364. [6] Iwai, K. (1976). Fundamental studies of fluid flow through a single fracture. Ph.D. Thesis. Univ. of California, Berkeley, USA. [7] Sheng, J. (2000). Studies on 3-D numerical simulation of fractured rock masses taking into account the coupling of seepage and stress and application. Ph.D. Thesis. Hohai University, Nanjing, P.R. China. [8] Zhu, Y. et al. (2002). Analysis of seepage control characteristics of the optional scheme of Three Gorges or RCC gravity dam on the Yangtze River. Journal of Hohai University. Vol.28(2), 1-6. [9] Zhu, Y., Gong, D., Zhang H. et al. (2003). Plane element simulation of fracture seepage for analyzing seepage in RCCD. Journal of Hydraulic Engineering, Vol 3, 63-68. [10] Zhu Y., Liu C., Li J. (2005). Study on seepage control methods and watertight and drainage measures in RCCDs. Guizhou Water Power, Vol. 19(3), 5-11. [11] Zhu Y. (1997). The computing method of equivalent nodal virtual discharge in Darcy's equation. Journal of Hohai University. Vol. 25(4), 105-108. [12] Zhu Y., Zhang, L., Pang Z. (1999). Study on permeability properties of RCC Dams and Longtan gravity RCCD (1). Hongshui River. Vol. 18(1), 2-8. [13] Zhu Y., Zhang L., Pang Z. (1999). Study on permeability properties of RCC Dams and Longtan gravity RCCD (2). Hongshui River. Vol. 18(2), 8-11. NAVODILA AVTORJEM Članki so objavljeni v angleškem jeziku s prevodom izvlečka v slovenski jezik. vsebina članka Članek naj bo napisan v naslednji obliki: - Naslov, ki primerno opisuje vsebino članka in ne presega 80 znakov. - Izvleček, ki naj bo skrajšana oblika članka in naj ne presega 250 besed. Izvleček mora vsebovati osnove, jedro in cilje raziskave, uporabljeno metodologijo dela, povzetek izidov in osnovne sklepe. - Uvod, v katerem naj bo pregled novejšega stanja in zadostne informacije za razumevanje ter pregled izidov dela, predstavljenih v članku. - Teorija. - Eksperimentalni del, ki naj vsebuje podatke o postavitvi preiskusa in metode, uporabljene pri pridobitvi izidov. - Izidi, ki naj bodo jasno prikazani, po potrebi v obliki slik in preglednic. - Razprava, v kateri naj bodo prikazane povezave in posplošitve, uporabljene za pridobitev izidov. Prikazana naj bo tudi pomembnost izidov in primerjava s poprej objavljenimi deli. - Sklepi, v katerih naj bo prikazan en ali več sklepov, ki izhajajo iz izidov in razprave. - Literatura, ki mora biti v besedilu oštevilčena zaporedno in označena z oglatimi oklepaji [1] ter na koncu članka zbrana v seznamu literature. oblika članka Besedilo naj bo pisano na listih formata A4, z dvojnim presledkom med vrstami in s 3.0 cm širokim robom, da je dovolj prostora za popravke lektorjev. Najbolje je, da pripravite besedilo v urejevalniku Microsoft Word. Hkrati dostavite odtis članka na papirju, vključno z vsemi slikami in preglednicami ter identično kopijo v elektronski obliki. Enačbe naj bodo v besedilu postavljene v ločene vrstice in na desnem robu označene s tekočo številko v okroglih oklepajih. enote in okrajšave V besedilu, preglednicah in slikah uporabljajte le standardne označbe in okrajšave SI. Simbole fizikalnih veličin v besedilu pišite poševno (npr. v, T itn.). Simbole enot, ki sestojijo iz črk, pa pokončno (npr. Pa, m itn.). Vse okrajšave naj bodo, ko se prvič pojavijo, izpisane v celoti. slike Slike morajo biti zaporedno oštevilčene in označene, v besedilu in podnaslovu, kot sl. 1, sl. 2 itn. Posnete naj bodo v kateremkoli od razširjenih formatov, npr. BMP, JPG, GIF. Za pripravo diagramov in risb priporočamo CDR format (CorelDraw), saj so slike v njem vektorske in jih lahko pri končni obdelavi preprosto povečujemo ali pomanjšujemo. Pri označevanju osi v diagramih, kadar je le mogoče, uporabite označbe veličin (npr. v, T). V diagramih z več krivuljami mora biti vsaka krivulja označena. Pomen oznake mora biti razložen v podnapisu slike. Za vse slike po fotografskih posnetkih je treba priložiti izvirne fotografije ali kakovostno narejen posnetek. preglednice Preglednice morajo biti zaporedno oštevilčene in označene, v besedilu in podnaslovu, kot preglednica 1, preglednica 2 itn. V preglednicah ne uporabljajte izpisanih imen veličin, ampak samo ustrezne simbole. K fizikalnim količinam, npr. t (pisano poševno), pripišite enote (pisano pokončno) v novo vrsto brez oklepajev. Vse opombe naj bodo označene z uporabo dvignjene številke1. seznam literature Vsa literatura mora biti navedena v seznamu na koncu članka v prikazani obliki po vrsti za revije, zbornike in knjige: [1] Feng, T. W. (2000). Fall-cone penetration and water content ralationship of clays. Geotechnique 50, No. 2, 181-187. [2] Ortolan, Ž. and Mihalinec, Z. (1998). Plasticity index-Indicator of shear strength and a major axis of geotechnical modelling. Proceedings of the Elev- enth Danube-European conference on soil mechanics and geotechnical engineering, Poreč, 25 -29 May 1998. [3] Toporišič, J. (1994). Slovenski pravopis. 2nd.ed., DZS, Ljubljana. podatki o avtorjih Članku priložite tudi podatke o avtorjih: imena, nazive, popolne poštne naslove, številke telefona in faksa, naslove elektronske pošte. Navedite kontaktno osebo. sprcjcm člankov in avtorske pravice Uredništvo si pridržuje pravico do odločanja o sprejemu članka za objavo, strokovno oceno mednarodnih recenzentov in morebitnem predlogu za krajšanje ali izpopolnitev ter terminološke in jezikovne korekture. Avtor mora predložiti pisno izjavo, da je besedilo njegovo izvirno delo in ni bilo v dani obliki še nikjer objavljeno. Z objavo preidejo avtorske pravice na revijo ACTA GEOTECHNICA SLOVENICA. Pri morebitnih kasnejših objavah mora biti AGS navedena kot vir. Rokopisi člankov ostanejo v arhivu AGS. Vsa nadaljnja pojasnila daje: Uredništvo ACTA GEOTECHNICA SLOVENICA Univerza v Mariboru Fakulteta za gradbeništvo Smetanova ulica 17 2000 Maribor Slovenija E-pošta: ags@uni-mb.si INSTRUCTIONS FOR AUTHORS The papers are published in English with a translation of the abstract into Slovene. format of the paper The paper should have the following structure: - A Title that adequately describes the content of the paper and should not exceed 80 characters; - An Abstract, which should be viewed as a mini version of the paper and should not exceed 250 words. 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Microsoft Word for Windows is the preferred format for submission. One hard copy, including all figures, tables and illustrations and an identical electronic version of the manuscript must be submitted simultaneously. Equations should be on a separate line in the main body of the text and marked on the right-hand side of the page with numbers in round brackets. units and abbreviations Only standard SI symbols and abbreviations should be used in the text, tables and figures. Symbols for physical quantities in the text should be written in Italics (e.g. v, T, etc.). Symbols for units that consist of letters should be in plain text (e.g. Pa, m, etc.). All abbreviations should be spelt out in full on first appearance. figures Figures must be cited in consecutive numerical order in the text and referred to in both the text and the caption as Fig. 1, Fig. 2, etc. Figures may be saved in any common format, e.g. BMP, JPG, GIF. 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Indicate the corresponding person. acceptance of papers and copyright The Editorial Committee of the Slovenian Geotechnical Review reserves the right to decide whether a paper is acceptable for publication, to obtain peer reviews for submitted papers, and if necessary, to require changes in the content, length or language. Authors must also enclose a written statement that the paper is original unpublished work, and not under consideration for publication elsewhere. On publication, copyright for the paper shall pass to the ACTA GEOTECHNICA SLOVENICA. The AGS must be stated as a source in all later publication. Papers will be kept in the archives of the AGS. tables Tables must be cited in consecutive numerical order in the text and referred to in both the text and the caption as Table 1, Table 2, etc. The use of names for quantities in tables should be avoided if possible: corresponding symbols are preferred. In addition to the physical quantity, e.g. t (in Italics), units (normal text), should be added on a new line without brackets. Any footnotes should be indicated by the use of the superscript1. For further information contact: Editorial Board ACTA GEOTECHNICA SLOVENICA University of Maribor Faculty of Civil Engineering Smetanova ulica 17 2000 Maribor Slovenia E-mail: ags@uni-mb.si list of references References should be collected at the end of the paper in the following styles for journals, proceedings and books, respectively: [1] Feng, T. W. (2000). Fall-cone penetration and water content ralationship of clays. Geotechnique 50, No. 2, 181-187. [2] Ortolan, Ž. and Mihalinec, Z. (1998). Plasticity index-Indicator of shear strength and a major axis of geotechnical modelling. Proceedings of the Eleventh Danube-European conference on soil mechanics and geotechnical engineering, Poreč, 25 -29 May 1998. NAMEN Revije Namen revije ACTA GEOTECHNICA SLOVENICA je objavljanje kakovostnih teoretičnih člankov z novih pomembnih področij geomehanike in geotehnike, ki bodo dolgoročno vplivali na temeljne in praktične vidike teh področij. ACTA GEOTECHNICA SLOVENICA objavlja članke s področij: mehanika zemljin in kamnin, inženirska geologija, okoljska geotehnika, geosintetika, geotehnične konstrukcije, numerične in analitične metode, računalniško modeliranje, optimizacija geotehničnih konstrukcij, terenske in laboratorijske preiskave. Revija redno izhaja dvakrat letno. aims and scope ACTA GEOTECHNICA SLOVENICA aims to play an important role in publishing high-quality, theoretical papers from important and emerging areas that will have a lasting impact on fundamental and practical aspects of geomechanics and geotechnical engineering. 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