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17-22 SEPTEMBER 2017


Contents Mede-ondersteuners

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ELGIP vision on reduction of geotechnical uncertainties for infrastructure J. Breedeveld / M. Woning

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Added value of advanced methods for New Large Diameter Sampler for soft soils seismic verification of existing www.pao.tudel . vertical cylindrical liquid storage Dr. ir. C. Zwanenburg / S. Luijendijk, M.Sc. ™ ™Leiderdorp Instruments BV, Leiderdorp ™Geomil Equipment BV, Moordrecht Associate Members Me mbe rs tanks and their foundations ™ ™Votquenne Foundations NV, Dadizele (B) ™ ™JLD Contracting BV, Edam F. Besseling MSc / M. Versluis MSc

™ ™Tjaden BV, Heerjansdam

A. Bougioukos MSc




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ELGIP vision on reduction of geotechnical uncertainties for infrastructure

Joost Breedeveld Deltares

Mike Woning Deltares

Abstract Transport infrastructure, including its connecting hubs that enable intermodal transport, are either built on or in the subsoil, and often use soil as construction material. Consequently, the subsoil plays a critical role in the complete life cycle of transport infrastructure. However, the natural formation and (sub)base materials that support the actual surf-

ace or structure are often very much overlooked as an important part of the infrastructure system, with occasionally disastrous consequences. This article summarizes the ELGIP vision on future transport infrastructure which aims at highly optimized, risk management-driven geotechnical (re)design, construction, maintenance and operation.

Figure 1 - One solution in “Forever Open Roads” vision by FEHRL

Introduction Transport infrastructure, including connecting hubs that enable intermodal transport, are either built on or in the subsoil, and often use soil as construction material. Consequently, the subsoil plays a critical role in transport infrastructure design, construction, maintenance and demolition. However, the natural formation and (sub)base materials that support the actual surface or structure (e.g. tunnel) are often very much overlooked as an important part of the infrastructure system, as can be seen in figure 1. With occasionally disastrous consequences… A strong group of 13 European research organisations in geotechnical engineering, the European Large Geotechnical Institutes Platform (ELGIP, see figure 2), aims to promote the profession internationally and has taken matters in its own hands. With over 2000 of professional staff, its members are committed to show that geotechnical engineering is essential in dealing with many pressing societal challenges associated with the built environment including transport infrastructure. Figure 2 - ELGIP (

ELGIP members representing Norway (Norwegian Geotechnical Institute), Sweden (Swedish Geotechnical Institute), the Czech Republic (Technical University of Prague) and the Ne-

therlands (Deltares) have formulated a vision document ‘Reduction of geotechnical uncertainties for transport infrastructure’. This vision describes the challenges concerning the use and complete life cycle of transport infrastructure, and explains how geotechnical engineering plays a prominent role in solving these. The ELGIP vision on future transport infrastructure aims at highly optimized, risk management-driven geotechnical (re)design, maintenance and operation. This article will highlight the main components of this vision document. The impact of geotechnical engineering First, two examples of transport infrastructure disasters are given, to underpin the importance


of the natural formation and supporting subsoil materials: •  On December 6th 2006, during the construction of a new part of road E6 in a quick clay area at Munkedal in Sweden, a landslide occurred affecting the old road (see figure 3). Several cars were drawn into the landslide. About 500m of road and 200m of the adjacent railway were destroyed. Fortunately, no one died. The costs for reconstruction alone were about €52 million. The landslide occurred due to incorrectly stored masses of subbase materials that triggered the slide [1]. •  On March 20th 2012 a retaining structure along motorway A13 in Austria between Innsbruck and Brenner to Schönberg sud-


denly collapsed. This 40 year old concrete structure was designed according to the standards that had to be met at that time. It was regularly inspected, but failed only weeks after the last inspection (see figure 4). The retaining wall failed extremely rapid due to a combination of unexpected loading (by water accumulation behind the wall due to exceptionally high snow melt), structural problems and brittle behavior. As a result a truck driver was killed. Also potential risk led to the control of other similar retaining walls, and after the evaluation some parts were reconstructed. To put this in perspective, in the years 20002006 the European Union (EU) invested €859

billion in its transport infrastructure [2], corresponding to €122 billion annually. Based on collected examples of similar disasters, it seems fair to assume that the failure costs equal at least 10% of the investment costs. Extrapolated to the EU, total failure costs may amount to €12.2 billion. And a conservative estimation of subsoil-related failures (about 1/3 of total failure costs) then amounts to about €4 billion annually for the EU. Hence, geotechnical engineering plays an important role in one of the greatest challenges of modern society: continuing to provide a safe, secure, efficient and affordable transportation network for people and goods. The resulting (geo)technical challenge is twofold:

Figure 3 - Road damaged by landslide at E6, Munkedal, Sweden

1.  New transport infrastructure and hubs need to be built in a more resilient, more durable and more affordable manner; 2. Existing transport infrastructure need to be maintained, retrofitted and repurposed to meet societal demands. Policy challenges As mentioned in policy documents from the European Commission (EC), transport is a key factor in modern economies [3]. Infrastructure is essential for the European quality of life, see figure 5, and vital for the EU’s competitiveness [4]. The required infrastructure network enables links between the different stages of production chains and allows service industries to reach their clients. Moreover, mobility is a significant employer in its own right. According to EC policy documents challenges for transport networks focus on the availability, affordability and sustainability of the infrastructure.

Source: SGI.

An infinitely available infrastructure network The 2001 EU Transport White Paper [3] also addresses the permanent contradiction between society demanding ever more mobility, and public opinion becoming increasingly intolerant of chronic delays and poor quality transport services. When transport systems are efficient, they provide economic and social opportunities: e.g. better accessibility to markets, employment and additional investments. Conversely, when transport systems are deficient in terms of capacity or reliability (and thus not available), they can have an economic cost such as reduced or missed opportunities and lower quality of life. Figure 4 - Schematics of retaining wall (a) and scene of failure (b)




Figure 5 - Infrastructure shapes mobility.

In that respect congestion is a major concern. Research has shown that in Utrecht, Manchester and Paris drivers spend more than 70 hours per year in road traffic jams [5]. And congestion costs Europe about 1% of gross domestic product (GDP) every year [6], which represented approximately €14.5 trillion in 2015 [7].

Source: Chennai sustainable transportation network

Unequal development of transport infrastructure between neighboring regions also has a negative influence on their interconnectivity. For example the Trans-European Transport Networks (TEN-T), which represent 800 km of key European corridors, have 9 north-south connections linking the continent, but only 4 east-west ones. And knowing that building a motorway, from planning to construction, can take up to 20 years, improvements of infrastructure networks have to be planned far ahead. With regard to infrastructure planning, in 2006 a list of EC policy actions [8] addressed the need to ensure a balanced approach to land-use planning.

Figure 6 - Trans-European Transport Network (TEN-T).

By decreasing uncertainties in the natural formation and of subsoil materials through innovations in geotechnical engineering, significant gains may be achieved for infrastructure availability. Better understanding of local subsoil behavior and soil-structure interaction enable more efficient and timely maintenance strategies and less disruptive maintenance techniques. Moreover, improvements in geotechnical risk management and monitoring, also during the infrastructure’s lifetime, will lead to less conservative observation-based design and construction. An affordable infrastructure network A well-performing transport network requires substantial resources. In 2011 the cost of EU infrastructure development to match transport demand has been estimated [9] at over €1.5 trillion for 2010-2030. And the completion of the TEN-T network (see figure 6) would require about €550 billion by 2020. Obviously, there is an increased pressure on public resources for infrastructure funding. User pay for the transport infrastructure network. However, not all costs related to the network are fully covered by the individual transport users (e.g. congestion, environmental damage and accidents). And the degree to which infrastructure costs are covered varies significantly both within and across modes. New approaches to funding and pricing of transport is required that reflect all costs of infrastructure. The affordability of transport infrastructure is clearly



Figure 7 - Wash-out of road and railway embankment at Ånn, Sweden.

Source: SGI

What geotechnical engineering has to offer… The ELGIP Vision Document illustrates that innovations in geotechnical engineering will have a significant positive impact on the availability, affordability and sustainability of transport infrastructure networks.

Source: Boston Globe

Figure 8 - Result of Boston’s Big Dig tunnel project.

linked to a reduction of its life cycle costs. This includes extending the life span of existing infrastructure and increasing its resilience, see figure 7. Additionally, new infrastructure needs to be steadfast to maintain its long term functionality under changing conditions. Knowing that at least one-third of infrastructure failure costs are subsoil-related, geotechnical innovations will have a significant impact on the affordability of infrastructure. These cover, amongst others, the development of reliable early warning systems for network parts vulnerable to hazards and methods to assess the condition of existing geotechnical structure (e.g. embankments, slopes). A sustainable infrastructure network Preferably, transport infrastructure invest-

Innovations in geotechnical engineering boost sustainable infrastructure through capacity increasement of (existing) transport infrastructure, while at the same time a lower energy demand (during construction), lower raw material inputs and a smaller spatial footprint are required. It enables a sustainable transport infrastructure network that reduces health and safety risks during natural disasters, accidents and unwanted events and supplies (geothermal) energy.

ments are planned to maximize positive impact on economic growth and minimize negative impact on the environment, see figure 8. The importance of sustainability is emphasized in Europe2020, the EU’s ten-year growth and jobs strategy launched in 2010. It addresses the shortcomings of our growth model and aims to create the conditions for smart, sustainable and inclusive growth. The Europe2020 Resource-efficient Europe flagship initiative supports the shift towards a resource-efficient, low-carbon economy to achieve sustainable growth. And in alignment with this strategy, the 2011 Transport White Paper of the EC adopted a roadmap of 40 concrete initiatives for a competitive and resource efficient transport system (e.g. dramatically reducing GHG emissions in transport by 2050).


The main difference between subsoil materials (e.g. sand, gravel) and other building materials (i.e. steel, concrete, and to a lesser extent timber) is that subsoil materials are a natural material with much larger spatial variability, determined both by the environment at the time of deposition and the following geological history (see figure 9). This is accompanied by a much larger variability of subsoil characteristics. As a conservative estimate, earth structures may show uncertainties about 50% in the final required specifications whereas timber, concrete and steel structures show uncertainties in the range of 3-20%. In addition, determination of soil parameters for geotechnical design is influenced by the type and extent of ground investigations, and the subsequent interpretation made by the geotechnical engineer. This results in a complex interaction that makes a reliable choice of geotechnical design parameters challenging. In this context, the geotechnical community currently handles inherent subsoil variability in two ways: • By increasing the extent of ground investigations in an attempt to model the subsoil with more detail. • By implementing an (over)-conservative design with the use of safety factors. This may lead to unnecessarily expensive and less sustainable design of geotechnical structures that demand too many natural resources. The ELGIP vision for sustainable, available and affordable transport infrastructure focusses, amongst others, on optimal observation-based geotechnical design. Leaner, less



Figuur 9 - Variability of soil profiles over short distances.

Source: Brady & Weil, 2002.

Figure 10 - Betuweroute railway line

conservative designs would result in substantial savings in construction costs and environmental impact, without affecting its stability and durability. In addition, the identification, assessment and prioritization of geotechnical risks for existing and newly built infrastructure will help to coordinate and economically apply the resources to minimize, monitor and control potential geotechnical hazards that could affect them. This is known as geotechnical risk management. Example The Waardse Alliance (the Netherlands) was related to the construction of part of the Betuweroute railway line, in which the subsoilrelated risks were fully shared by client and contractor. The client asked for effective and cost-efficient solutions for building in challenging soft soil conditions. As part of the project, systematic instrumentation (based on risk management) was used for monitoring the con-

struction process, aiming at achieving savings and increasing in the efficiency of it. The sustainability impact of the project dealt with the optimization of land and resources (i.e. sand) used, resulting in minimized construction time and barriers for the surroundings. The availability impact of the project was reflected in the completion of the project within the expected time frame, enabling the operations to start on time. This directly links to the ‘availability’ ambitions. Due to the innovative approach, a positive financial project result of €25 million was achieved. Conclusion The ELGIP vision document shows that the application of risk management in geotechnical engineering in general, and monitoring or continuous control of subsoil conditions in particular, has great potential in leading to significant advantages for our society. Currently, the lack

Tabel 1 - ELGIP objectives for future risk management-driven transport infrastructure (re) design, maintenance and operation





Guiding objective

Failure frequency, e.g. due to man-made and natural disasters


Delay duration due to infrastructure repair, maintenance, reconstruction


Fatalities and severe injuries due to man-made and natural disasters


Travel time of persons / goods


Total Cost of Ownership


Land use for infrastructure network


Use of raw materials


Use of secondary materials



of sufficient research and innovation prevents us from using this potential, gaining the advantages for society. Table 1 below shows the ELGIP objectives for the future of highly optimized, risk management-driven geotechnical (re)design, construction, maintenance and operation for infrastructure networks. Reference [1] “Skadekostnader i byggprocessen– En litteratur genomgång”, SGI Varia 642, Linkoping 2012; [2] Steer Davies Gleave, 2009, "Ex Post Evaluation of Cohesion Policy Programmes 20002006, Work Package 5A: Transport", First Intermediate Report; [3]  COM(2001) 370 final, WHITE PAPER EU transport policy for 2010: time to decide, Brussels, 12.9.2001; [4]  COM(2008) 433 final, Greening Transport, Brussels, 8.7.2008; [5] INRIX European National Traffic Scorecard 2010; [6] Transport 2050: The major challenges, the key measures, EC memo/11/197, Brussels, 28.3.2011; [7] see gross-domestic-product-gdp-in-the-european-union-eu/; [8] COM(2006) 314 final, Keep Europe moving Mid-term review of the European Commission’s 2001 Transport White Paper, Brussels, 22.06.2006; [9] EC calculations based on TENtec Information System and the Impact Assessment accompanying the White Paper, SEC(2011) 358;


F. Besseling MSc Team leader Earthquake engineering & Structural dynamics at Witteveen+Bos

Added value of advanced methods for seismic verification of existing

M. Versluis MSc Structural engineer Hydraulic structures at Witteveen+Bos

vertical cylindrical liquid storage tanks and their foundations

A. Bougioukos MSc Structural engineer Earthquake engineering & Structural dynamics at Witteveen+Bos

Hydrodynamic pressures When a liquid storage structure is dynamically excited, hydrodynamic pressures occur on walls, bottom and possibly on the roof. This first has been elaborated for dams by Westergaard [1]. Later, i.a. Housner [2] has developed similar solutions for (rigid) liquid storage tanks in which a distinction is made between impulsive and convective pressures. i.a. Veletsos and Yang [3] have introduced a third term which takes into account the flexibility of the structure. These three terms of the total (horizontal) hydrodynamic pressure are explained in equation (1).



In the these expressions a cylindrical coordinate system (r,z,θ) is used with the origin at the centre of the tank bottom and the positive z-axis in the direction of the fluid. Rigid and flexible impulsive pressures The impulsive pressure is the result of the fluid mass which moves with the container wall. A distinction is made between the mass that moves with the ground acceleration (rigid) and the mass that moves with the acceleration difference between ground and wall acceleration (flexible). The rigid impulsive pressure distribution has the shape of a parabola, the flexible impulsive pressure distribution depends on the mode shapes of the structure. Figure 1 shows normalized impulsive pressure distributions calculated with equation (2) which can be found in EN 1998-4 and [3].


In case of a rigid tank (Ďˆ(z) = 1), the expression for the flexible impulsive pressure will result in the expressions for rigid impulsive pressure. For liquid storage containers the ratio H/R (fluid height / radius) is the most important geometrical parameter that determines the seismic load level. This is because the opposing walls influence each other. In the case of navigation locks and dams it can be shown that when the cham-

ber or reservoir length is more than 4 times the water depth, the pressure is independent of the geometry besides the water depth [1,6].

shows the pressure distribution over the height of the container which is calculated according to equation (3).

Convective pressures The convective pressure results from the sloshing of the fluid in standing waves. Figure 2

Because the fundamental eigen-period of the convective mass is in most cases for groundsupported structures much larger than the fun-




Earthquakes cause (hydro)dynamic actions on (steel) liquid storage tanks. In static conditions the tank shell of a vertical cylindrical liquid storage tank is mainly subjected to circumferential (ring) forces. Earthquake excitation results an additional horizontal load component to the tank structure, disturbing the axi-symmetric stress state. This can cause damage if tanks are not properly designed for this specific type of load. Designs of new to be built tanks are often easily adjusted at some specific aspects in order to increase seismic resistance. Typically, simplified design codes that cover these important structural aspects are used for design. More advanced verification/design analysis specifically have

significant added value for existing storage tanks not designed to resist seismic loads or for design optimizations in general. This paper focuses on these advanced seismic verification methods and relates them to the behaviour of tanks under seismic response and relevant failure mechanisms. In addition this paper focuses on specific issues that are relevant in relation to seismic design of tank foundations and liquefaction hazard and soil-structure interaction effects. Special focus is put on the shortcomings and limitations of simplified methods and the benefits that more advanced additional verifications can have for clients that develop or exploit tank storage facilities.

Figure 1 - Horizontal impulsive pressure distributions on tank wall for H/R = 0.8 and θ = 0

It is important to realize that the vertical excitation of the earthquake also causes hydrodynamic pressures. Although these are axisymmetric (and therefore do not result in an overturning moment), they cause an increase/ decrease of the total internal pressure. Because the internal pressures influences the buckling resistance in steel tanks, the vertical excitation direction should be included in the analysis.

Figure 2 - Convective impulsive pressure distributions on tank wall for H/R = 0.8 and θ = 0

Failure mechanisms The following sections describe some of the most critical failure modes observed for liquid storage tanks subjected to earthquake motions. It is not intended to provide a complete list of possible failure modes here. Instead, we address some failure modes of specific interest in relation to analysis methods described in other sections of this paper.

ring only the fundamental convective mode. The eigen-periods follow equation 4.

Overturning In case of horizontal earthquake excitation, mass inertia of the tank structure and the liquid product of hydrodynamic pressures result in global overturning. Anchored and unanchored tanks can be distinguished. For anchored tanks specific anchor elements mount the tank superstructure to the foundation. These anchors should be designed to resist overturning action. For unanchored tanks overturning stability is provided by the weight of the fluid on the bottom plate and more specifically on the annular ring. This annular ring is a strengthened ring plate that should have sufficient width to accommodate plastic rotations and limit uplift of the tank. Post earthquake observations of tank failures in some cases shown poor performance for tanks without a properly designed or constructed annular ring configuration. Moreover, meridional (axial) compression stresses at the compressed side of unanchored tanks increase in case of uplift at the tension side. This should be accounted for in the design.

damental eigen-period of the impulsive mass, dynamic coupling between the modes can be neglected. In literature the convective pressure is therefore calculated based on the geometry (H/R) and independent of the flexibility of the fluid container. In case of elevated tanks the fundamental periods can be much closer to each other and dynamic coupling can only be disregarded if the fundamental periods are at least a factor 2.5 separated [7]. For ground-supported tanks of common H/R ratios the convective pressure contributions from higher order modes are negligible and can be sufficed with conside-


The compressibility of the fluid can also result in hydrodynamic pressures, but this is mainly relevant for high head dams and not for liquid storage tanks of common dimensions [1,6]. The resultant of the hydrodynamic pressure on the walls and tank bottom causes an overturning moment and a base shear.



Photo: J. Skogh (a,c) / (a,b,c)

Figure 3 - Typical buckling modes of tank

a) Elastic local buckling due to Livermore earthquake, 1979

b) Elastic-plastic local buckling (elephant-foot) due to the San Fernando earthquake, 1971

Shell buckling Overturning moment action results in increasing meridional membrane stresses. Two main dominant failure mechanisms can be observed, namely the elastic shell buckling (diamond shape buckling mode) and the elastic-plastic shell buckling (elephant’s foot buckling shape). In addition, in the area of the shell perpendicular to the direction of loading, shear buckling can occur. The three shell buckling modes are illustrated by figure 3. The internal hydrostatic and hydrodynamic pressures form an important parameter for shell buckling behaviour. Compared to empty tanks, increasing internal fluid pressure initially stabilizes the tank shell. However, with increasing internal pressure at some point the threshold level of allowable meridional stress drops rapidly as the hoop stress reaches the Von Mises yield stress. This is illustrated by figure 4, in which the horizontal axis shows a dimensionless coefficient of internal pressure (hoop stress versus yield stress) and the vertical axis shows the elastic ιxpe and elastic-plastic ιxpp pressurized reduction factors from EN 1993-1-6. The same principle holds for shear buckling. Internal pressures stabilizes the shell and when the hoop stress is 30% of the yield stress or larger, shear buckling does not occur [8]. In

c) Shear buckling due to Livermore earthquake, 1979

practice this means that lower shell courses of filled tanks are generally not susceptible to shear buckling tanks.

Figure 4 - Schematic influence of tensile hoop stress on the meridional buckling stress [9]

Annular ring failure The annular ring forms a strengthened segment of the bottom plate of a tank. Other portions of the bottom plate are typically constructed of steel plates with less thickness and lower weld capacities. Any plastic rotations should concentrate in the annular ring in order to prevent failures due to insufficient rotation capacity of the bottom plate. The width and welding detailing of the annular ring therefore are essential for the seismic resistance of unanchored tanks. Foundation failure Seismic foundation failure of storage tanks on shallow pad or plate foundations can relate to overturning resistance, liquefaction effects, or a combination of both. When during a design phase of a tank facility potentially liquefiable deposits are encountered, often piled foundations are constructed or liquefaction mitigation measures are taken during construction (e.g. dynamic compaction or stone columns). However, these measures have high consequences in terms of costs and impact on operations for existing tanks that would need measures to achieve sufficient safety against liquefaction. In this perspective detailed investigation of tank


foundation stability in relation to liquefaction is worthwhile. Analysis and design methods Simplified methods In Europe, the Eurocodes EN 1993-4-2 and EN 1998-4 are the main standards available for the (seismic) design and verification of liquid storage tanks. Specific rules for design and manufac-



Figure 5 - Fundamental mode of tank-fluid system

turing of welded steel vertical cylindrical tanks can be found in EN 14015. In addition, buckling of steel shell structures is covered by EN 19931-6. Annex A of EN 1998-4 and annex G of EN 14015 (based on the American API 650) give recommendations and design graphs/tables to calculate the overturning moment and base shear force with a simple, quasi-static approach of a MDOF system. Based on the H/R ratio, the impulsive and convective modal masses, as well as their acting heights, can be calculated. The sum of the impulsive and convective mass is the total fluid mass. The spectral acceleration of these modes follows from the response spectrum ordinates at the calculated periods. With this simplified method no distinction is made between the rigid and flexible impulsive mass. Instead, it is assumed that all impulsive mass acts rigid (the parabolic shape), but with the spectral acceleration corresponding to the fundamental eigen-period of the combined tank-fluid system instead of the ground acceleration. For the considered tanks (0,5 < H/R < 1,0), the resulting base shear and overturning moment from this simplified method match quite well with the results obtained with a modal analysis where each of the three hydrodynamic components are summed up. The same conclusion is drawn in [5], which also gives an elaborate explanation about these simplified methods. There are also drawbacks of these preliminary design methods: • The simplified methods do not always include the seismic vertically induced pressure, which influences buckling verifications for steel tanks; •  Shear buckling of steel tanks is not discussed; •   Influence of soil-structure interaction is usually not taken into account; •  Although the overall overturning moment may be correct, the flexibility of the tank wall can cause higher pressure at the top shell courses compared to an assumed rigid pressure distribution (as shown in figure 1). Designing the wall thickness of the upper shell courses with respect to earthquakes proportional to the lower shell course thickness and hydrostatic pressure, as mentioned in annex G4.4 of EN 14015, should therefore be treated with caution.

Modal analysis and FSI effects Besides the simplified method, also a FEM modal analysis can be performed, which can be incorporated in more elaborate evaluation methods like prescribed by EN 1998-4. This part of Eurocode series addresses coupling between hydrodynamic pressure distributions, eigen-periods and mode shapes. Considering FEM modal analysis, two methods can be identified: • Using added fluid mass in a structural model for eigen-value analysis. • Using a full fluid-structure interaction (FSI) model for eigen-value analysis. In the first method the following steps need to be performed: 1. Assume a mode shape for the tank walls; 2. Calculate the added mass distribution according to e.g. Veletsos and Yang [3]; 3. Include this added mass with the structural FEM model and perform modal analysis. 4. Extract the mode shape and repeat steps 2 through 4 until convergence is reached. The above procedure can be performed with most FEM software packages. Figure 5 shows the fundamental mode shape for the tank-fluid system. The second method approaches the problem by analysing the coupled system of the fluid and the structure; the so-called fluid-structure interaction. The problem is solved through the finite element method in which three types of elements are considered simultaneously; the structure, the fluid and the interface elements. In the structure, the discretization is given in the


following form: (5) where MS, CS, and KS are the mass, the damping and the stiffness matrices of the structure respectively, whereas u is the unknown displacement of the structure. The vector fI represents the interface forces due to interaction between the solid and the fluid, while the vector fSext represents the external load which acts on the structure. In the case of the fluid, the unknown quantity is the pressure variable p and the interface forces are considered for the coupling with the structure. At the interface (with the structure) boundary of the fluid the following equation holds: (6) where, ρF is the fluid density and nF the outward normal to the fluid domain. Considering continuity between the normal displacements of the fluid and the structure with the condition üF = üs, the last equation (6) becomes: (7) The discretization in the fluid is given in the following form: (8) Taking into account that in the element level the


Figure 6 - left: modelling of fluid-structure interface elements; right: modelling of boundary condition at the surface of the fluid (potential head).

contributions of the interface force is fIe=-ReT pe and rIe=ρF Re üe, the coupled system of equations can be obtained:

on shallow foundation introduces an additional degree of freedom in the system which results in an increase of the fundamental period and

 (9) This system can be solved either in the frequency or in the time domain depending on the type of the external force. Subsequently, an eigenvalue analysis is performed in order to identify the dominant eigen-modes. Modal analysis and SSI effects Codes and guidelines for seismic design differentiate two main categories of soil-structure interaction (SSI) effects, being inertial and kinematic interaction. For vertical cylindrical steel storage tanks, inertial interaction is most important. The non-rigid based condition for tanks

a potential increase of system damping. The change in system configuration can be visualized by considering the equivalent mechanical spring model of the tank with a rigid foundation on top of an another equivalent mechanical spring-dashpot model of the soil consisting out of translation and rocking springs and dashpots (illustrated by figure 7). Aiming at correct prediction of period elongation and increased damping the main challenge is to properly quantify the SSI springs and dashpots. Design codes and guidelines typically speFigure 7 - Mechanical model SSI principle.

cify SSI coefficients as function of foundation radius, soil shear modulus and soil Poisson’s ratio, with a frequency dependent multiplier [10]. Although this concept may appear quite simple and straightforward, the actual implementation has a few complicating facets. First of all, there may be a scale effect for moderate to large storage tanks on grade. Moreover, rigid slab assumption is not valid for unanchored tanks or tanks on pad foundation which have a relatively flexible bottom configuration. Rigid global rocking behaviour of the tank-foundation pad therefore can be unrealistic. Moreover, the simplified expressions include a single soil shear modulus, for which a constant value is assumed following perfect homogeneous half space conditions. In reality for layered soils shear modulus is not constant and moreover effective shear modulus due to modulus reduction is shear strain dependent and therefore seismic load dependent. The latter implies a non-constant shear modulus over the time frame of an earthquake and a varying effective shear modulus among different zones below and adjacent to the tank. In order to address these issues a combined approach is suggested in which code based expressions supplemented with finite element analysis from the basis of determination of SSI effects. An example for a 60 m diameter tank on layered alluvial deposits is shown below. The first figure illustrates measure shear wave velocity Vs (SCPTU) and correlated (from CPT) Vs and small-strain shear modulus G0 depth profiles. This information forms the basis for prediction of SSI constants with code based expressions. It can be read that within the influence depth of the tank the effective equivalent small strain shear modulus is approximately 75 MPa in this case.




Figure 8 - Example of SCPT measured Vs-depth profile and Vsdepth and G0-depth profiles form CPT + correlations

ling stress of a specific structure can be found, including boundary conditions (e.g. stiffening girders), varying diameter/thickness ratios or asymmetric (hydrodynamic) loading. But also for more standard tanks a higher critical buckling stress can be found compared to the classical expression: (10) The eigen-values are dependent of the internal pressure and thus on hydrodynamic loading. Figure 10 shows the first relevant eigen-value for shell buckling for a 60 m diameter tank at a maximum hoop stress of 69% of the yield stress. Based on LBA it was substantiated that the actual critical buckling stress of the specific tank is about 20% higher compared to the basic idealized solutions following from LA stress design.

Subsequently, figure 9 shows vertical deformation contours of a tank axi-symmetric model due to a 10% foundation load increase and similarly deformation contours for the tank subject to overturning loading. From these FEM calculations the effective global SSI behaviour can be better understood and static SSI constants can be derived. In this specific case it turned out that the code based formula needed adjusted input values in order to match both horizontal, vertical and rocking response stiffness. With the reported SSI constants the fundamental period of the system for the specific tank under consideration in this example elongated from approximately 0.37 s to 0.46 s.

Linear bifurcation analysis In order to determine the local buckling resistance of the shell several types of analyses can be performed, ranging from purely linear analytical expressions (LA stress design) up to full geometrically and physically non-linear analysis with imperfections (GMNIA). Another linear method is the FEM linear bifurcation analysis (LBA), which basically finds the critical Euler buckling force (lowest eigen-value) for the specific shell. In order to find the characteristic buckling stress, the knockdown factor from stress design are applied to the elastic buckling stress from the LBA. These knockdown factors include non-linear effects like imperfections and plasticity. Benefits of a LBA are that the elastic buck-

Pushover analysis The basic method for performing post-elastic seismic analysis of structural capacity is a nonlinear pushover analysis (NLPO). By increasing a load factor on the impulsive load the pushover curve can be calculated. This method can also be used for storage tanks to substantiate the sequencing of the relevant failure modes for a tank. Relevant nonlinearities for unanchored steel vertical cylindrical liquid storage tanks are: •  Foundation overturning and sliding bearing capacities; • Uplift (compression-only spring supports); •  Plastic material behaviour such that plastic hinges can occur; • Local buckling of the shell. One may differentiate ductile mechanism nonlinearities from non-ductile (or brittle) mecha-

Figure 9 - Vertical deformation contours for a tank subject to seismic differential foundation pressures (left: vertical motion; right: is rocking motion)



Figure 10 - Eigen-vector from LBA for a 60 m diameter tank .

Figure 11 - Pushover response spectrum analysis result in ADRS format.

A three-step approach is used to evaluate the seismic tank foundation integrity including liquefaction potential. The first step follows from the evaluation of hydrodynamic pressures and global overturning response of the tank, based on which the seismic foundation loads can be calculated. Secondly, empirical methods that predict liquefaction triggering potential predict the level of potential excess pore pressure generation as a function of seismic hazard, ground characteristic and tank configuration (e.g. [11]). In these assessments both effects of increased isotropic stress states below the tank and high shear stress zones next to the tank are accounted for. Then as a final step, a 3D finite element model including tank superstructure, pad foundation and soil stratigraphy is developed. In the layered soil stratigraphy liquefaction effects can be accounted for by explicit modelling of excess pore pressure generation or by means of equivalent reduction of effective friction angle. With this model the (reduced) seismic tank foundation stability can be assessed. This three-step approach overcomes the need for an effective stress time-dependent finite element model to evaluate liquefaction risks. Although such effective stress models are nowadays rapidly being developed we note here that their implementation currently needs really extensive calibration. One of the reasons for this statement is the impact of initial stress state, which can severely corrupt the numerical performance of such constitutive models [12].

nism nonlinearities. Ductile mechanisms, like uplift and - to some extend - plastic material behaviour, limit the seismic capacity but accommodate nonlinear response can be obeyed to design a tank structure for lower demand spectra. Pushover response spectrum analysis is the tool used for this purpose. As a first step, a nonlinear numerical model of the tank structure is developed. Then a unit load vector representing the impulsive pressure component is calculated. This unit load vector subsequently is applied to the nonlinear numerical model of the tank by gradually increasing its amplitude by scaling the load vector with some scalar. The pushover load range needs to extend well beyond the design seismic demand. The nonlinear pushover curve that is calculated from this analysis then can be coupled in an acceleration-displacement response spectrum

(ADRS) framework to the seismic demand and equivalent nonlinear system damping as illustrated by figure 11. Seismic design of tank foundations Tank shallow foundations are characterized by large differential stress levels below and next to the tank. As a result, the zone just next to the tank foundation edge shows high initial static shear stress levels in the soil. Depending on the relative density of the soil this can increase or decrease significantly the liquefaction triggering potential of the soil. This is illustrated by figure 12, where on the left hand side shows the static shear stress coefficient for cyclic stress liquefaction triggering analysis according to [11] and on the right hand side shows the impact for a specific tank configuration in terms of calculated excess pore pressure ratio.


Conclusions and recommendations The present paper illustrates the added value of integration of simplified and advanced techniques and methods for seismic design/verification of liquid storage tanks. Although simplified code based methods are able to capture the main elements important in seismic response and seismic resistant design, we observe that advanced analysis give much more insight and can be very helpful for improved evaluations. This is of specific interest for tanks not specifically designed for seismic action or tanks in areas subject to changes in seismic code regulations. Integrating advanced techniques in the verification procedure in this case enhances the quality of verifications, gives a more complete assessment of potentially critical elements and helps to optimize the design of any required seismic retrofitting. References [1].  Westergaard, H.M., â&#x20AC;&#x2DC;Water Pressures on Dams During Earthquakesâ&#x20AC;&#x2122;, Transactions,



American Society of Civil Engineers 98 (1933), nr. 1835, New York (USA), p. 418472. [2]. Housner, G.W., ‘Earthquake Pressures on Fluid Containers’, Eight Technical Report under Office of Naval Research, California Institute of Technology, Pasadena (USA) 1954. [3].  Veletsos, A.S., Yang, J.Y., ‘Dynamics of Fixed-Base Liquid-storage Tanks’, Proceedings of U.S.-Japan Seminar on Earthquake Engineering Research with Emphasis on Lifeline Systems, Tokyo (JP) 1976, p. 317-341. [4].   Scharf, K., ‘Beiträge zur Erfassung des Verhaltens von erdbebenerregten, oberirdischen Tankbauwerken’, Fortschritt-Berichte VDI, Reihe 4. Bauingenieurwesen, nr. 97 (1990), VDI Verlag, Düsseldorf (D).

[5].   Malhotra, P.K., Wenk, T., Wieland, M., ‘Simple Procedure for Seismic Analysis of Liquid- Storage Tanks’, Structural Engineering International 10 (2000), nr. 3, p. 197-201(5). [6]. Versluis, M., (2010), Hydrodynamic pressures on large lock structures, Master thesis TU Delft and Witteveen+Bos, 2010. [7].  IITK-GSDMA Guideline for Seismic Design of Liquid Storage Tanks - Provisions with Commentary and Explanatory Examples (2007), National Information Center of Earthquake Engineering, Indian Institute of Technology Kanpur. [8]. Design Recommendation for Storage Tanks and Their Supports With Emphasis on Seismic Design, Architectural Institute of Japan 2010, Tokyo (JP) 2010. [9].  Rotter, J.M., Schmidt, H., Buckling of Steel

Shells - European Design Recommendations, (5th edition), ECCS 125, ECCS, Brussel (BE) 2013. [10]. Gazetas, G., Analysis of Machine Foundation Vibrations: State-of-the-Art. Soil Dynamics and Earthquake Engineering. 1983, 2(1), 2-43. [11]. Idriss, I.M., Boulanger, R.W., Soil liquefaction during earthquakes, Monograph EERI MNO-12, Earthquake Engineering Research Institute, 2008. [12]. Elsäcker, W.A., Evaluation of seismic induced liquefaction and related effects on dynamic behaviour of anchored quay walls - using UBC3D-PLM constitutive model, Master thesis TU Delft and Witteveen+Bos, 2016.

Figure 12 - Left: example of static shear stress ratio near a tank foundation; right: corresponding concentrating excess pore pressure ratios calculated based on [11]

Figure 13 - Changing foundation failure modes as a result of liquefaction effects (figure shows shear strain contours plots close to failure)



Material Point Method and Applications in Geotechnical Engineering

Alexander Rohe Deltares

Mario Martinelli Deltares

Introduction A vast number of computational methods is being developed to simulate large deformation problems involving soil-water-structure interaction. Here, the material point method (MPM) is used that was developed to simulate large deformations in history-dependent materials. It combines the advantages of mesh-based and point-based approaches: mesh distortion is eliminated and history is stored in material points. Generally speaking, MPM is an advancement of the finite element method (FEM) where the continuum body is represented by a set of Lagrangian points, so-called material points (MPs). The MPs are moving through an Eulerian background mesh. The MPs carry all physical properties of the continuum such as stresses, strains, density, momentum, material parameters and other state parameters, whereas the background mesh is used to solve the balance equations without storing any permanent information.

computational costs and stability caused by material points crossing element boundaries. An overview of the historical development and applications of MPM can be found in several contributions of the proceedings of the first international conference on the material point method (Procedia Engineering, 2017). They provide a state-of-the-art overview of currently active groups developing MPM and recent advancements in applications for geotechnical and hydraulic engineering.

volume of the MP can change enabling material compression or extension. The second frame is the computational background mesh which is equivalent to a conventional finite element mesh. It is constructed to cover the full domain of the problem, i.e. also areas where material is expected to move into during computation. The discretised momentum balance equations are solved at the nodes of this computational mesh, whereas mass conservation and constitutive equations are solved at the MPs.

The material point method is implemented into the Anura3D MPM Software of the Anura3D MPM Research Community ( This is a collaboration of the Universitat Politècnica de Catalunya (UPC Barcelona), University of California Berkeley, University of Cambridge UK, Delft University of Technology, Deltares, Università degli Studi di Padova, Technische Universität Hamburg-Harburg (TUHH) and associated partner Università degli Studi di Salerno.

Additional benefit of MPM is the continuum description of the material for which well-known constitutive relations to describe the material’s stress-strain behaviour can be applied. Furthermore, the application of engineering boundary conditions is easily facilitated, i.e. stresses and displacements or their rates can directly be applied on any boundary or material point. Shortcomings of MPM are its mesh-dependency which is inherent to any finite element formulation,

Concept of Material Point Method (MPM) In MPM the computational domain is spatially discretised in two frames. First, the continuum body is divided into a set of material points (MPs). Each MP represents an initially defined volume of the domain with associated mass. One of the basic and most important features of MPM is that the mass of each MP strictly remains constant which implies that mass conservation is automatically satisfied, whilst the

The MPM algorithm for a single calculation step is illustrated in Figure 1. At the beginning of each step, the components of the momentum balance equations are defined by mapping information from the MPs to the nodes of the computational mesh by means of interpolation or shape functions (Fig. 1a). The equations of motion are solved for the primary unknown variables, i.e. the nodal accelerations (Fig. 1b). These nodal values are used to update acceleration, velocity and position of MPs, as well as to compute strains and stresses at the MPs (Fig. 1c). The assignment of MPs to elements is updated after mesh adjustment (Fig. 1d). For technical details the reader is referred to e.g. Rohe & Martinelli (2017). Soil, in general, is a mixture of three constituents (solid, liquid and gas) which interact with each other determining the mechanical and hydraulic response of the material. However, taking rigorously into account these interactions

Adapted from MPM Research Community, 2017

Figure 1 - Computational scheme of MPM.




The concept of the material point method (MPM) is briefly introduced for the numerical modelling of large deformations and soil-water-structure interaction for applications in geotechnical and hydraulic engineering. The formulation of MPM for a fully-coupled two-phase porous material in contact and interaction with free surface water is used. The concept of multi-

ple sets of material points is introduced which enables modelling of free surface water, groundwater flow, liquefaction and erosion problems. The numerical framework is implemented in the Anura3D MPM Software. Two application examples are presented, i.e. slope failure due to infiltration and due to seepage flow.

Adapted from Yerro et al., 2015

Figure 2 - Concept of multiple constituents in MPM approaches depending on number of phases and number of material point sets as used in Anura3D.

Figure 3 - Example of the 2-point MPM schematisation of a partly saturated slope in contact and interaction with free surface water.

In the 2-phase 1-point approach each material point describes a representative volume element of fully saturated soil, carrying the information of both phases, liquid and solid together. While the material points are attached to the solid skeleton giving a Lagrangian description of its movement, an Eulerian approach with respect to the solid represents the motion of the liquid phase. Typical applications are consolidation, pile installation (jacking, driving, vibrating) with generation and dissipation of excess pore pressures, CPTs in partially drained conditions and failure of saturated slopes due to infiltration or loading. A full formulation of the 2-phase 1-point approach can be found in MPM Research Community (2017).

of material point sets to describe each phase. The grey shaded combinations are currently available in the MPM implementation of Anura3D and are described below.

In the 3-phase 1-point approach, the soil is understood as a material composed of three distinct constituents, i.e. solid, liquid and gas. The solid phase constitutes the soil skeleton of the media while liquid and gas phases fill the voids. All phases are combined in a single material point and balance and momentum equations are formulated and numerically solved as above. No independent motion of the water and air is expected in the applications envisaged. Typical applications are rainfall and drought effects in slope failure and collapse analyses for unsaturated soils, collapse behaviour of low-density soils and unrestrained swelling of expansive clays. The full formulation of the 3-phase 1-point approach can be found in Yerro et al. (2015).

Modelling of dry soil, pure liquid and saturated soil in drained and undrained conditions are the

In the 2-phase 2-point approach the solid-liquid mixture is modelled using two distinct sets of

Adapted from Martinelli, 2016 may be in many cases unnecessarily complicated, computationally expensive, and even unfeasible for engineering applications. In Figure 2 the concept of multiple constituents as used in the MPM formulation of Anura3D is shown. The rows represent the number of phases of the continuum, the columns represent the number

fields of applicability of the 1-phase 1-point formulation. In this case only one set of material points is needed to carry all required information of the material. Typical applications are undrained slope failure, collapse of dry granular materials, silo discharge problems, water reservoirs, shallow foundations in undrained conditions, undrained cone penetration tests (CPTs) and submerged slope failure in highly permeable sandy soils. The full formulation of the 1-phase 1-point approach can be found in e.g. MPM Research Community (2017).



Adapted from Soga et al., 2016

Figure 4 - Anura3D results for modelling slope failure due to infiltration. Evolution of a) equivalent plastic shear strain; b) excess pore water pressures [kPa ]; for time t* in whole slope. Evolution of c) total displacements [m]; d) velocity [m/s]; in four points of the slope.

material points, one for each constituent respectively. Figure 3 shows an example of a slope which is characterised by two materials: soil (dry and saturated parts) and free surface water. The soil is a porous material composed of two constituents, the solid skeleton and groundwater respectively, whereas the free surface water is a pure liquid. In general, the water can flow out of the soil body into the water reservoir or vice-versa, and interacts with the solid skeleton through drag forces. The soil can behave as a solid porous material with liquid in its pores (solidlike response) or as a liquefied material in which soil grains float in the liquid (liquid-like response). The material can also change between these states. Application examples are transient seepage flow through a porous material or the free-fall of a poro-elastic body under water (e.g. dropping of geocontainers). Advanced examples involving state transition are the free-fall of a coarse-grained soil under water, collapse of a submerged sand column (liquefaction and breaching), collapse of a slope due to seepage flow with subsequent sediment transport, internal instability (suffusion), piping, erosion problems, slurry and debris flows including separation. The full formulation of the 2-phase 2-point approach can be found in Martinelli (2016).

Application 1: Slope failure due to infiltration The 2-phase 1-point formulation is used to analyse the failure of a slope due to water infiltration below the dike. The Selborne slope failure experiment (Cooper, 1998) is simulated in which progressive failure in overconsolidated clays was investigated in nearly fully saturated conditions. The instability process is illustrated in Figure 4. The evolution of equivalent plastic strain and excess pore water pressures is shown at five normalised times t*=t/Tfailure. Note that the plastic shear strain scale varies in order to capture the results properly. The calculated total displacements and velocity of four points along the slope are shown. The progressive nature of the slope failure can be observed. The slope remains stable until t*=1. When the failure mechanism develops those material points located above the shear band (1, 2, 3) accelerate quickly and a peak velocity is attained at t*=1.1. At t*=1.2 the slope tends to stabilise. Point 4 remains motionless and point 3 at the toe moves only a small distance. The development of failure starting at the toe and crest of the slope, geometry of the failure surface, measured displacements of the slope and the time history of pore pres-


sures are successfully simulated using the 2-phase 1-point MPM formulation in Anura3D. It is shown that MPM can adequately simulate the problem, and results compare well with field measurements of the Selborne experiment. Further details can be found in Soga et al. (2016). Application 2: Slope failure due to seepage flow The 2-phase 2-point formulation is used to analyse the failure of a dike slope due to seepage flow. During the calculation the water head in the upstream reservoir (left) drops and water flows through the slope towards the downstream reservoir. Figure 5 shows the soil porosity at the location of the solid material points together with the position of the liquid material points at four time steps. The porosity varies between 40% (initial value) and 100% (pure water) but is only shown in the range below the maximum soil porosity of 52%. During the simulation the pore pressure at the toe increases and the drag forces by the flowing water move the solid material points rightwards. Plastic shear strains are induced in the soil and the porosity increases accordingly, which is due to the positive dilatancy angle assumed in the elastoplastic Mohr-Coulomb soil constitutive model.



At the beginning (t=6s), the porosity increases only at the toe whereas, as the simulation proceeds (t=8s), the change in porosity propagates also along the shear plane. At the toe the porosity increases until it exceeds the limit porosity and fluidises. As the time progresses (t=14s), the fluidised zone increases, moving upwards towards the crest, and the fluidised material located at the toe is moving into the downstream reservoir. A large increase in porosity is also observed along the failure surface (t=23s), where some solid material points exhibit a fluid-like behaviour. This evidence is related to the basic constitutive model which has a constant dilatancy angle independent of the plastic strains. Therefore, along the failure surface where most of the shear strains are concentrated, the response of the constitutive model produces a progressive increase in volume which leads to fluidisation of the soil. In future analyses a more advanced critical state constitutive model is necessary for modelling more realistic soil behaviour. Conclusion The material point method (MPM) is introduced and used for simulating large deformation problems, and there seems an opportunity to utilise this technique to assess the risk of damage after progressing failure. MPM was selected as

preferred choice, primarily because: (a) the implementation is intuitive for users of FEM; (b) it can incorporate advanced history-dependent soil constitutive models; (c) its application of boundary conditions is more straightforward than other mesh-free methods owing to the presence of the background grid; and (d) fully coupled behaviour of soil, pore water and free surface water can be included in the formulation. The example applications demonstrate the potential opportunities of MPM for large-deformation soil-water-structure interaction analysis to predict both pre- and post-failure behaviour. The authors would like to thank Jurjen van Deen (Deltares) for his assistance in editing this paper. References -  Cooper M.R., Bromhead E.N., Petley D.J., Grant D.I. (1998): The Selborne cutting stability experiment. Géotechnique 48(1):83-101. DOI: 10.1680/geot.1998.48.1.83. -  Martinelli M. (2016): Soil-water interaction with material point method - Double-point formulation. Deltares report, MPM-DREDGE (PIAP-GA-2012-324522). - Martinelli M., Rohe A., Soga K. (2017): Modeling dike failure using the material point method. Procedia Engineering 175:341-348. DOI: 10.1016/j.proeng.2017.01.042. - MPM Research Community (2017): Anura3D

MPM Software, Scientific Manual, v2017.1. Eds.: Martinelli M. - Procedia Engineering (2017): Proceedings of the 1st International Conference on the Material Point Method, MPM2017, 10-13 January 2017, Delft, The Netherlands. Eds.: Rohe A., Soga K., Teunissen H., Zuada Coelho B., Volume 175, Pages 1-372. - Rohe A., Martinelli M. (2017): Material point method and applications in geotechnical engineering. In: Proceedings of the Workshop on Numerical Methods in Geotechnics (NMG2017). Hamburg University of Technology. - Soga K., Alonso E., Yerro A., Kumar K., Bandara S. (2016): Trends in large-deformation analysis of landslide mass movements with particular emphasis on the material point method. Géotechnique 66(3):248-273. DOI: 10.1680/ jgeot.15.LM.005. - Yerro A., Alonso E.E., Pinyol N.M. (2015): The material point method for unsaturated soils. Géotechnique, 65(3):201-217. DOI: 10.1680/ geot.14.P.163.

Adapted from Martinelli et al., 2017

Figure 5 - Anura3D results for modelling slope failure due to seepage flow. Position of liquid material points (blue coloured) and soil porosity of solid material points (colour legend) is shown for several time steps.



New Large Diameter Sampler for soft soils

Introduction Since 2017 a new large diameter sampler is operable. The sampler is developed for retrieving high quality soft soil samples with a diameter of 0.4 m and a height of 0.5 or 1.0 m. This article discusses the reasons for developing the sampler, its use and preliminary results. One of the issues when dealing with peat is its fibrous nature. In conventional geotechnical laboratory testing, the test results are usually interpreted by an approach based on continuum mechanics. Such an approach is only allowed when the dimensions of individual particles are smaller than the dimensions of the tested specimen. As a rule of thumb it is assumed that the largest dimensions of the individual particles are at least 10-times smaller than the smallest dimension of the specimen. Most conventional laboratory tests are axial symmetric with a diameter of 3.8, 5.0 or 6.5 cm. This is in the same order of magnitude as the length of peat fibres. As a consequence size effects are to be expected when testing peat samples. The results of laboratory testing often are used in geotechnical design of water retaining structures and road improvements. When the parameters are more accurate and reliable, this can save money by optimizing the geotechnical designs.

Dr. ir. Cor Zwanenburg Deltares, The Netherlands

Sophia Luijendijk, M.Sc. Deltares, The Netherlands

0.40 m and a height of 0.50 or 1.00 m. Besides the use for retrieving large diameter samples, the sampler can also be used as an alternative for block sampling of soft clays. Than the large diameter sample is trimmed in the laboratory to the required dimensions.

Figure 1 - Components of the DLDS.

Start of design The design of the large diameter sampler is started with a literature study. The study aims for finding the state of the art in understanding sample disturbance. The available literature can be roughly divided into two groups. One group discusses numerical studies on sampling and sample disturbance (a.o. Mohsen et al 1987, Clayton et al 1998). This group studies the disturbance of an idealized material behaviour by an idealized sampling method. A second group discusses field experiences (a.o. Long et al 2009, Santagata et al 2006, Tanaka et al. 1996). In this group often the total disturbance, due to sampling, transportation and laboratory handling is




top cap Sample tube knives (inside) cutting shoe

Figure 2 - Detail of DLDS: Ring with cutting knives, red rings are for transportation purpose only.

In earlier studies (Zwanenburg & Van (2015)) sample size effects are found for failure and post failure behaviour. Failure of the fibres, either by slippage or rupture occur after some displacement. Although the strain levels in conventional and large sized samples are the same, the actual displacement is different, leading to different failure behaviour. In order to test this behaviour from undisturbed peat samples large dimensions are required. To facilitate large sized testing a large sized sampler is required. This paper discusses the development of Deltares Large Diameter Sampler, DLDS, which is capable of retrieving samples with a diameter of




A new large diameter sampler, DLDS, has been developed and is operable since 2017. This sampler can retrieve undisturbed samples with 0.4 m in diameter and 0.5 or 1.0 m in height. The large sampler has two purposes. First it facilitates laboratory testing on large volumes; for fibrous peats it is shown that size effects play a role in laboratory testing. A large triaxial device and a large direct simple shear device have been developed to study

Figure 3 - Detail of DLDS: Sample tube.

protection tube

sample tube knives Cutting shoe

discussed by comparing laboratory test results. Literature specific on sampling peat is limited (a.o. Helenelund et al 1972, Long 2006, Mathijssen et al 2008). For sampling peats, a sharp cutting edge, to make clear cuts through the fibres is important. Dragging fibres down with the sampler should be avoided. Lunne et al (1997, 2006) presents a sample disturbance index for clays based on the void ratio change by reloading the sample to field stress conditions. A low void ratio change during reloading indicates little disturbance. It is questionable if this index can also be used for peats. Due to the high compressibility of peat in combination due to high permeability the sample easily compresses during sampling and a low void ratio change during reloading to field stress conditions might not necessarily indicate little sample disturbance. Sampler equipment From the beginning the purpose was to design a sampler for sampling large diameter peat and soft clays in high quality environment. Without lateral support the sample might bulge when bringing the sample from sampling depth to ground level. For this reason it was decided to design the new sampler as a tube sampler. During lifting of the sample, trans-

the behaviour of large samples in comparison to conventional sized samples. The sampler is developed to facilitate these large scale tests. Second, the sampler can be used to retrieve high quality samples in soft, organic clay and peat for conventional laboratory testing as an alternative for block sampling. In establishing the quality of the samples a comparison is made to samples retrieved with the well-known Sherbrooke sampler.

portation and further handling of the sample, the tube provides the required lateral support. The sampler is also designed as a down-hole sampler, meaning that first a boring is realized to the desired sampling depth. Collapse of the borehole walls is prevented by a casing. From this basic idea, a down-hole tube sampler, a design is made containing the following items, see Figure 1: - Cutting shoe -  Knives, to cut the sample after the required depth is reached, see Figure 2. - The tube, to collect the sample, see Figure 3. - Top cap which includes a suction valve and the connection to the plunger, see Figure 4 -  The plunger that pushes the sampler down, see Figure 4. The samples are taken from a pre-drilled borehole. At two elevations along the plunger, three struts, six in total, can be pushed against the casing, fixing the position of the sampler. Sample handling Due to its dimensions and weight, the large samples are not easy to handle by manpower without disturbing the sample. Therefore, a special procedure and required tools are developed for further sample handling. After the sampler is pushed into the soil, the suction valve is closed. The struts, for vertical fix, are released and the sampler is lifted to ground level. Experience shows that for soft soils the cutting shoe is empty, no soil is stuck inside. The sample rests on the knives. In case of an empty cutting shoe a pedestal is constructed. The sampler and cutting shoe are placed over the pedestal, such that after releasing the suction valve the soft sample is supported by the pedestal. Then the plunger and top cap are removed and the top of the sample is inspected. The recovery ratio is measured and the space between the top of the sample and top of the sample tube is filled with soft light weight material. The top of the sampling tube is carefully, water tight, closed by a lid.


The sampler is lifted again and carefully turned upside down. To make lifting possible, lifting rings can be screwed onto the sides of the sampler. Next the cutting shoe and ring with knives are removed. A lid is placed at the, new, top. Finally, the sample can be transported to the laboratory for further handling and testing. Application in the field To test the sampler, samples are retrieved from the Uitdam test site, near Amsterdam, located in the north of the Netherlands, Wiertsema (2016). There a large, rather homogenous peat layer is easily accessible, to retrieve samples. At the Uitdam test site a series of field trials are conducted to test the operational shear strength of peat, for details see Zwanenburg & Jardine, (2015). At the site also Sherbrooke samples were retrieved, which gives the option to compare the quality of the samples obtained by both methods. The subsoil at the Uitdam site consists of a 4 to 5 m thick peat layer which overlays a 4 to 5 m thick clay deposit followed by a Pleistocene sand layer. The samples were taken from the peat layer at a depth, top sample, of 1.39 m below ground level. The peat comprises mainly Phragmites, with sedge and sphagnum inclusions, with minor vegetal decomposition. The peat is characterized with a von Post classification of H2 to H3, indicating minor decomposition, a water content ranging between 650 and 1250 %, an organic content ranging from 75 to 92% and a particle density of 1.53 Âą 1.6% Mg/m3. The undrained shear strength, su ranges between 5 to 10 kPa. After careful inspection some improvements are made, but there was no indication that fibres were dragged through sample during sampling and the fibres at the outer radius of the sample were clear cut by the sampler. There are more sites in the Netherlands where the large diameter sampler is applied for research purposes. The results are not available yet, but will be published soon. Laboratory testing To further study sample quality a series of oe-


Figure 4 - Detail of DLDS: Plunger and top cap.

dometer tests is conducted on specimen taken from the DLDS samples and compared to results from tests on specimen from samples taken by the Sherbrooke sampler. Figure 5 shows preliminary results comparing classical oedometer tests on 4 specimen from DLDS samples and 2 specimen from Sherbrooke samples. The oedometer curves, see figure below, show good agreement. The differences between the different sampled specimens seem small. Table 1 gives further details on water content, w, dry density, γd, pre-consolidation stress, σ’vy, stiffness parameter CR and ratio between reloading and normally consolidated stiffness, CR/RR. Although it should be noted that the number of tests are too small to draw final conclusions, there seems to be a difference between the DLDS and the Sherbrooke samples. The DLDS specimens have a higher water content, lower density a lower normally consolidated stiffness and larger CR/RR ratio. There seems no clear difference in pre-consolidation stress. Differences in w, γd and ratio CR/RR can be explained by differences in sampling techniques in which some compression of the peat samples might have occurred. During sampling and bringing the Sherbrooke samples to ground level, the samples are not supported and water from the large pores can leave the samples easily. Some of the Sherbrooke samples showed some deformation, bulging or bending directly after sampling. However, the differences in CR and the lack of difference in σ’vy cannot be explained by sample disturbance. Alternatively, heterogeneity in the peat layer might also explain the differences between the test results. It should be noted that the samples were taken at close distance, centre to centre distance between the borings is 8.6 m. From visual inspection there was no indication for geological or biological differences in the peat layer. More test results are needed for final conclusions. Conclusions The new large diameter sampler, for soft soils, is designed, built and applied successfully in the field. Retrieved peat samples from this large diameter sampler are tested in the laboratory. The preliminary results show that the peat samples have less sample disturbance (so slightly higher quality) than the retrieved samples with conventional sample techniques (with a diameter of about 63mm) and the Sherbrooke sampler. Especially for peat the large diameter sampler seems to give high quality samples.




Figure 5 - Comparison between classical oedometer test results for tests on samples retrieved by large diameter sampler (DLDS) and Sherbrooke (SHER).

Tabel 1 - Comparison between the DLDS and Sherbrooke samples. test ID

w [%]

γd [kN/m3]

σ’vy [kN/m2]

CR [-]

CR/RR [-]





































The purpose of the sampler is twofold. First it aims to retrieve samples for large volume testing. Second, the sampler forms an alternative for block sampling, giving less sample disturbance. The sampler is applied in the Netherlands at the Uitdam test field, but could be applied on more sites. The sampler is well suited for sampling soft soils, such as peat and soft clay. A large triaxial device and a large direct simple shear device have been developed to study the behaviour of large samples in comparison to conventional sized samples. The sampler is developed to facilitate these large scale tests. Further research and comparison of the Uitdam peat sample in the large direct simple shear device is being planned for the coming period. References - Clayton C.R.I. Siddique A., Hopper R.J. (1998) Effects of design on tube sampling disturbance – numerical and analytical investigations Géotechnique 48 no 6. p 847-867 -  Henelund K.V., Lindqist L-O, Sundman C.

- Mathijssen F.A.J.M., Boylan N., Long M., Molenkamp F. (2008) Sample disturbance of organic soils geotechnical and geophysical site characterization, Huang & Mayne (eds) Taylor & Francis group, London, ISBN 978-0-41546936-4 - Mohsen M., Baligh M., Azzouz S., Chin C-T (1987) Disturbances due to “ideal” tube sampling journal of geotechnical engineering vol 113 no 7 p739-757 -  Santagate M., Sinfield J.V., Germaine J.T. (2006) Laboratory simulation of field sampling: comparison with ideal sampling and field data Journal of geotechnical and geoenvironmental engineering 132 no 3 p351-362 - Tanaka H. Sharma P., Tsuchida T, Tanaka M. (1996) comparative study on sample quality using several types of samplers Soils and Foundations 36 no 2 p57-68 -  Wiertsema (2016) Sherbrooke Sampler vs DLDS < watch?v=LHslW_iRAPg > (June 30th , 2017) - Zwanenburg C. Jardine R.J. (2015) Laboratory, in situ and full-scale load tests to assess flood embankment stability on peat Géotechnique 65 no 4 p 309-326 - Zwanenburg C., Van M.A., (2015) Comparison between conventional and large scale triaxial tests on peat 15th Pan American Conference on Soil Mechanics and Geotechnical engineering (Manzanal D., Sfriso A.O. eds) Buenos Aires, IOS Amsterdam

(1972) influence of sampling disturbance on the engineering properties of peat samples Proc of the international peat congress, Otaniemi - Long M. (2006) Use of a downhole block sampler for very soft organic soils Geotechnical testing journal 29 no 5 p 426-438 - Long M., El Hadj N., Hagberg K. (2009) Quality of conventional fixed piston samples of Norwegian soft clay Journal of geotechnical and geoenvironmental engineering 135 no 2 p 185198 - Lunne T., Berre T., Strandvik S. (1997) Sample disturbance effects in soft low plastic Norwegian clay in: Recent developments in soil and pavement mechanics Almeida (ed.) Balkema Rotterdam ISBN 90 54 10 885 1 - Lunne T., Berre T., Andersen K.H., Strandvik S., Sjursen M. (2006) Effects of sample disturbance and consolidation procedures on measured shear strength of soft marine Norwegian clays Canadian Geotechnical Journal 43 p 726-750



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