Advanced Dynamic Soil Testing – Introducing the New Marum...

Advanced Dynamic Soil Testing – Introducing the New Marum Dynamic Triaxial Testing Device

S. Kreiter, T. Moerz, M. Strasser, M. Lange, W. Schunn, B.F. Schlue, D. Otto, and A. Kopf

Abstract Soil mechanical and submarine mass-movement initiation studies often use static and quasi-static approaches to determine the strength of soils against external mechanical stresses. However, many natural processes pose time vari-ant stresses on soils, and hence exert key roles for submarine slope stability and submarine mass-movement initiation. Prominent examples are earthquake-, wind-, wave- and current-forces and alternating man-made loading on offshore construc-tions. Most soils show a weaker response to periodic loading – making dynamic and cyclic loading experiments mandatory for offshore natural hazard and risk assessment. Dynamic and cyclic triaxial testing are essential in liquefaction studies of granular soils and creep investigations of cohesive and granular sediments. So far, competing setups are used with mechanical spindles, pneumatic actuators or full hydraulic drives.

The new MARUM dynamic triaxial testing device (DTTD) unit is addressing this increasing demand by enabling a wide range of test configurations. At its core it contains an ultra fast, hydraulically-driven ±20 kN cylinder and a 5 kHz real-time controller. This enables up to ±0.5 mm strokes at up to 50 Hz. Advantages to com-mercial systems are (1) the high flexibility in test setup, (2) the possibility to feed arbitrary signals derived from in situ measurements, and (3) full system access to all controls to expand and adjust the system abilities on the hard and software level. Applications so far include cyclic creep studies for offshore wind farms, liquefac-tion experiments on artificial sand–clay mixtures and studies on the behavior of submerged soils under dynamic stress conditions to evaluate slope stability and submarine landslide initiation.

S. Kreiter, T. Moerz (�), M. Strasser, M. Lange, W. Schunn, D. Otto, and A. KopfMARUM – Center for Marine Environmental Sciences and Faculty of Geosciences, University Bremen, Leobener Str. 28334 Bremen, Germanye-mail: [email protected]

B.F. SchlueARCADIS Consult GmbH, Europaplatz, 3, 64293 Darmstadt, Germany

D.C. Mosher et al. (eds.), Submarine Mass Movements and Their Consequences, 31Advances in Natural and Technological Hazards Research, Vol 28,© Springer Science + Business Media B.V. 2010

32 S. Kreiter et al.

1 Introduction

The response of continental margins to time variant stresses has become an important element in offshore risk assessment. Dynamic and cyclic loading exerts a key role in submarine slope stability and tsunamigenic submarine landslide initiation (Sultan et al. 2004; Biscontin et al. 2004). Cyclic loading is important for long-term securing of offshore installation such as wind energy plants (Lesny and Hinz 2006; Lesny and Wiemann 2005). The time variant stresses from seismic or storm loading may either affect the natural system directly or be transmitted through manmade constructions. In the last few years, large research efforts have been undertaken to understand the influence of rapid and transient stress changes on slope stability in the marine realm (Sultan et al. 2004, 2008; Biscontin et al. 2004, Biscontin and Pestana 2006; Azizian and Popescu 2006, Stegmann et al. 2007; and many others). Under cyclic loading, the sediment behavior is influenced by the intensity and dura-tion of the cyclic stress. The soil properties relevant for the cyclic and dynamic behavior are the stress history, density or consistency respectively, sensitivity (for cohesive soils), degree of saturation, grain size distribution, and fine content (Kramer 1996). Cyclic stresses may lead to plastic deformation and to degradation in the stress–strain and strength properties of the sediment, which may result in the initiation of sediment mobilization. Additionally sands may show total liquefaction (Ishihara 1985; Pestana et al. 2000).

To simulate and study the wide variety of cyclic and episodic natural processes in the laboratory a testing system with maximum flexibility is required. Various dynamic shear systems have been developed over the previous decades. The majority of them are limited in their performance with regard to the frequency of the load applied. During recent years, results of undrained monotonic and cyclic ring shear tests on saturated cohesionless soils have been reported in the literature by Trandafir and Sassa (2004, 2005a, b), which are limited to 5 Hz. Research cyclic triaxial test systems at e.g. the Kyoto University or at the University of Duisburg-Essen work with pneumatic actuators or a combination of servo motor and pneumatic actuaors to apply the time variant loads up to approximately 1 Hz. The latest generation of commercially available dynamic triaxial systems are driven by electro-hydraulic servo actuators at 0.5 kHz, limiting their maximum frequency to 10 Hz or less (see www.geocomp.com and www.oce.uri.edu). In order to overcome these limitations, to avoid the disadvantage of restricted modification access using commercial sys-tems, and to address sediment deformation with earthquake loads, which frequency spectra contain relevant energies up to 20 Hz, the MARUM dynamic triaxial testing device (DTTD) project was initiated.

2 MARUM Dynamic Triaxial Testing Device

The MARUM DTTD project was started in 2005 when the German Science foun-dation granted the second phase of the DFG Research Center MARUM (http://www.marum.de). The system was developed in house from components with the

Advanced Dynamic Soil Testing 33

initial guidance of the industry partner Hansa-Flex Hydraulik GmbH – a company specialized in hydraulic system design and setup (http://www.hansa-flex.com).

The test equipment compasses a servo-driven hydraulic cylinder with a free configurable real time controller, a hydraulic power unit, displacement-, load- and pore pressure transducers, a load frame, a pneumatically controlled confining and backpressure unit, and a control station with a user interface to initiate the testing scenarios (Fig. 1). At the heart of the system is a servofloat® quality hydraulic testing cylinder built by Herbert–Hähnchen GmbH & Co. KG. This cylinder contains a patented annular gap seal and is characterized by stick-slip-free cylinder move-ments and ultra low-friction at all piston speeds. The test cylinder operates at up to 32 MPa (25 kN) and the maximum piston speed is 4 m/s. The test cylinder is physi-cally controlled by a 10 l/min D765 series servo MOOG valve directly mounted on to the cylinder via a voltage interface. The system performance is enhanced by two 1 l accumulator tanks (orange bulbs on the photo of Fig. 1) that prevent insufficient flux at the feed and the return side of the MOOG valve.

The oil cooling unit of the (11 kW) hydraulic power unit proved to be valuable in providing a constant oil viscosity and cylinder response. The unit delivers 28 MPa with a maximum flux of 15 l/min, allowing maximum oil flux at full servo valve opening which leads to frequencies up to 50 Hz and distance amplitude in access of ±0.5 mm.

The confining- and backpressures are provided via three independent servo-controlled valves (AirCom GmbH) of the PQ type with a pressure range from 0 to 2 MPa.

Fig. 1 Photograph and schematic drawing of the MARUM dynamic triaxial testing device


The valves are equipped with separate pressure transducers that report the pressure back via a 0–10 V interface. The heavy load frame is built around two parallel guid-ing rods that allow height adjustment for different experimental cells of up to 40 cm width. The load frame is mounted to massive concrete walls at the basement of the MARUM building, internal vibrations are suppressed via rubber shock-absorbers.

The real time data acquisition, processing and control unit (ADwin-light-16) comprises a SHARC™ digital signal processor from Analog Devices with a 40 MHz 32-bit FPU. Load and distance information is analog to digital converted and digit-ally processed by low phase digital filtering send to the control PC and to the PID (proportional, integral and differential) servo algorithm (Fig. 2). The resulting con-trol signal is connected to the MOOG valve via an analog output of the ADwin system. All calculations are repeated every 0.2 ms resulting in a 5 kHz feedback loop and two orders of magnitude oversampling at maximum frequency of 50 Hz.

The distance information is currently provided via (1) an internal inductive dis-placement sensor (WLH 250-K-0.5%, MESSOTRON) within the hydraulic cylinder and (2) an external laser optical triangulation system from MICRO-EPSILON.

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Fig. 2 Flowcharts of the force and displacement control processes in the real-time 5 kHz circuit


The internal inductive displacement control proved to be too temperature-sensitive and its resolution did not allow us to quantify sample deformations with the desired precision. However, the inductive sensor output is still used for safety reasons where distance control over the full stroke length is required. The external class II laser sensor ILD 1700 (Fig. 1) with a measuring range of 100 mm has a dynamic resolution of 0.003 mm (Fig. 3) at 2.5 kHz sampling frequency – accurate and fast enough to track the strain of stiff materials under dynamic loads of up to 50 Hz.

Load information is acquired via high-precision force transducers from GTM (Gassmann Testing and Metrology GmbH) with a range of 20 kN and 0.002% full scale accuracy. The GTM load transducers are directly mounted to the hydraulic cylinder rod (Fig. 1). Two ultra-low noise carrier frequency HBM amplifiers are used to collect the data of the force- and differential pore pressure transducers (Validyne) prior to analog/digital conversion. The input signal is logged and proc-essed at 5 kHz by the ADwin unit.

Process control and data storage is accomplished with the scenario planning & control station, a PC-based system running a Labview (National Instruments™) routine developed in-house. The control station coordinates the real-time ADwin unit with the back- and side pressure unit. This main user interface allows for the performance of user-defined scenarios in force-controlled and displacement-controlled modes (Fig. 2).

3 Performance Examples

To evaluate the actual status (capability, compliance, etc.) of the MARUM DTTD, various performance tests have been conducted on materials of variable stiffness. In Figs. 3 and 4, experimental results on the stiffest end member, an aluminum cylinder placed inside a standard triaxial testing cell, are presented. It can be seen that at 1 Hz and 60 kPa cyclic loading, the input function and system response are in excellent agreement. At 10 Hz and 60 kPa dynamic loading, a phase offset and slight waveform distortion can be observed.

Fig. 3 System response using an aluminum test cylinder at 1 Hz left and at 10 Hz right. Red solid lines: predefined stress; blue dashed lines: measured stress, green dotted lines: sample deformation


Frequencies of up to 10 Hz are well reproduced on stiff material (Fig. 4). However, the phase offset increases linearly for higher frequencies. Phase and amplitude offsets of the system response at higher frequencies or softer materials are tackled in an ongoing development effort to automatically modify the PID settings using a dynamic load adaptation (Fig. 2) This procedure will be especially valuable in experiments where soil properties are subject to change during complex runs.

The arbitrary signal function offers the possibility for testing soils with self-generated or recorded time – load signals to analyze the soil response to complex, but realistic multi-frequency dynamic loading scenarios under laboratory conditions. The performance example given here simulates a ground-motion record of the 1989 Loma Prieta earthquake (Mw 7.1) of a stratum 2 m subbottom. The time versus stress input function was derived using a procedure proposed by Seed and Idriss (1971) based on the empirical stress reduction factor from Iwasaki et al. (1978):

The shear stress function t(t) on a potential horizontal failure plane at depth z of a soil column moving horizontally as a rigid body according to a horizontal accel-eration function a(t) (i.e. the signal recorded by strong-motion seismometers) can be calculated using Newton’s laws of motion. In reality the soil column is internally deformable and the actual shear stress is less than that estimated from the above described rigid body model. This discrepancy is empirically corrected by the stress reduction coefficient r

d. The shear stress evolution t(t) is therefore defined by

d.

a(t)(t) r

g z

⋅ γτ =

⋅ (1)

where g is the unit weight and g is the gravitational acceleration. On the basis of a variety of soil-response data sets, Iwasaki et al. (1978), proposed r

d = 1 − 0.015

z/m. This allows for the assessment of shear stress history of a soil at a given depth from seismometer acceleration recordings.

Figure 5 shows the input arbitrary signal function and the measured system response measured on a soft foam test specimen (Young’s modulus 235 kPa).

Fig. 4 Summary of the MARUM dynamic triaxial device response on an aluminum test cylinder at different frequencies


The satisfactory reproduction of the input signal confirms that the DTTD system is able to adequately reproduce the ground acceleration of a given earthquake. Hence, any other arbitrary load signal derived from storm wave and tidal loading, traffic load (railways and motorways), foundations (bridges, dams, offshore platforms, etc.), or data from in situ measurements can be simulated.

4 Data Examples

4.1 Liquefaction of Sand

When initially loose, saturated granular sediment is exposed to strong dynamic or cyclic stresses (e.g. during an earthquake), the grains may reorganize to take up less space viz. the sediment may contract. Under undrained conditions the tendency toward pore water release reduces the effective stress by increasing the excess pore-water pressure. High cyclic stress amplitudes on loose granular soils may cause a drop in the effective stress to zero, which implies that the stress is taken up by the fluid rather than the particles: the soil is then liquefied (e.g. Ishihara 1985; Ishihara and Tsukamoto 2004).

Two variables are often used in cyclic failure assessment: the cyclic stress ratio (CSR) and the cyclic resistance ratio (CRR) (Kramer 1996). The cyclic stress ratio CSR is defined as s

d,cyc/2s′

3,i, where s

d,cyc is the cyclic deviator stress and s′

3,i

is the initial effective minor principal stress at the start of the cyclic loading. Hence, the CSR is a measure of the cyclic load. The CRR is the CSR for a given number of cycles at failure. Hence, the CRR is a measure for the capacity of a sediment

Fig. 5 Load time history: red solid line: arbitrary signal input function derived from acceleration time history of strong motion record from the Gilroy station #1, Channel 1 – at 090° for the 1989 Loma Prieta earthquake (Mw 7.1; 11 km from the rupture area) at Gilroy (CA), data from the Centre for Engineering Strong Motion Data (CESMD) http://strongmotioncenter.org. Blue dashed line actual load measured on the load sensor


layer to resist liquefaction and stress softening (Sultan et al. 2004). The ratio between both values is the factor of cyclic safety.

The CRR of natural quartz-rich dune sand has been determined with the new MARUM DTTD for various configurations. The sand was initially cleaned of organic material and sieved to a grain size ranging from 0.1 to 0.45 mm. The void ratio was e

max = 0.59 in its densest state and e

min = 0.96 in its loosest state. The sand

was prepared dry in the cell and subsequently saturated by 500 ml flow of deaired water through the sample. Multiple samples were then back-pressured up to 500 kPa until full saturation (B-values > 95%) was reached and consolidated to the desired initial effective minor principal stress s′

3,i. Then the samples were

subjected to cyclic loading by applying load cycles with different CSR values and frequencies.

Figure 6 presents an example result of such a cyclic triaxial test with dense sand at a density index I

D = 0.87, corresponding to an initial void ratio of e

0 = 0.64, a

B-value = 0.96 at 500 kPa backpressure, consolidated s’ 30 kPa, equivalent to an in-situ depth of ∼5 m. The CSR was set to 0.3 at 1 Hz, which is typical for strong earthquakes (see Seed and Idriss 1971, and Eq. 1). The normalized pore pressure (pore pressure/total stress) increases continuously with increasing numbers of cycles, reaching 1 after ∼17½ cycles (dashed line Fig. 6). Deformation starts after 15¾ cycles, when normalized pore pressure values are around 0.8 (solid line in Fig. 6). The negative strain reveals a dilatational behavior of the sand during lique-faction. As proposed by Kramer (1996) liquefaction of a soil element is considered to occur at excess pore pressures equal to 90% of the mean total stress (dotted vertical line in Fig. 6). Hence, the cyclic resistance of this sand under the given circumstances is equivalent to 16½ cycles, resulting in a CRR of 0.3 at 16½ cycles.

Fig. 6 Cyclic triaxial test of an artificial sand sample. Top graph: axial strain; bottom graph: pore pressure evolution. Vertical lines mark different liquefaction definitions: Solid line is onset of axial deformation; dotted line is pore pressure reaches 90% of total stress; dashed line is pore pressure reaches 100% of total stress


4.2 Cyclic Creep in Clays

The behavior of cohesive sediments under cyclic load is fundamentally different to the usual cyclic response of granular media. Commonly stiffness and shear strength of clayey sediments degrade under cyclic loads. Figure 7a illustrates cyclic stress-induced deformation behavior of slightly overconsolidated silty clay, tested with every 10,000 cycles increased cycles cyclic stress ratios. At low CSRs the clay deforms incrementally during the first 1,000 cycles, followed by a shake down in deformation. At higher cyclic stress levels the deformation does not cease com-pletely, and the clay deforms at lower and lower rates not likely leading to failure (stabilization). Under even higher CSRs the plastic deformations of clays may accumulate with ongoing cyclic loading and lead to progressive failure; however such a behavior has not been observed here for the CSR tested (Fig. 7a). Shake down, stabilization and progressive failure are the three most common behavior types of cohesive sediments (Mallikarjuna 1992; Lesny and Richwien 2004).

But cohesive soils show a variety of creep deformation responses. Figure 7b illustrates a CSR independent creep of strongly overconsolidated plastic clay. Here the number of stress cycles seems to be more important for the deformation behavior than the incrementally increased CSR. So far it is unclear whether (1) a critical consolidation ratio exists, (2) if a certain type of mineralogy is required for CSR independent creep behavior, and (3) if such a behavior could be described by a cyclic deformation rate, describing the material deformation by a rate to be multiplied by the number of stress cycles, not regarding the CSR. These are open questions currently addressed in an ongoing study.

5 Summary and Conclusion

The increasing awareness of the importance of cyclic and dynamic loading in hazard initiation stimulated worldwide efforts to build adequate testing devices. In this paper, the setup of the MARUM DTTD is presented and the crucial components

Fig. 7 Cyclic triaxial test results of samples from a North Sea wind farm area. Cyclic frequency is 5 Hz, cyclic s

1 is stepwise increased every ~10,000 cycles. Blue solid line: axial strain; green

dotted line: CSR. (a) Slightly overconsolidated silty clay, s3,i

is 200 kPa. (b) Highly overconsolidated clay, s

3,i is 100 kPa


and control functions are discussed in the context of commercial systems. The full dynamic capabilities of servo-valve controlled hydraulic systems, however, can only be exploited if the force and displacement transducers have excellent accuracy and dynamics. The programmable digital signal processor of the DTTD allows flexible system operation modes. A series of performance tests have been per-formed to benchmark the system’s capabilities. Frequencies up to 10 Hz can be used with confidence for soil testing. Higher frequencies up to 50 Hz are easily reached by the DTTD system, but lead to large phase offsets. The linear trend in acceleration suggests that the hydraulic system is capable of even higher dynamics. An earthquake equivalent stress signal is reproduced confirming that the MARUM DTTD is capable of accurately simulating multi-frequent arbitrary signals. Test carried out using standardized sand samples resulted in high quality strain and pore pressure data that allow a reliable CRR determination. Tests on highly overconsoli-dated clays revealed critical cyclic behavior of CSR-independent creep, stimulating further in depth studies. Cyclic creep, cyclic softening and liquefaction tests are now routinely conducted at the MARUM in the context of various ongoing projects, such as (1) the design of offshore wind power plants, and (2) slope stability analysis and submarine landslides studies (e.g. 1979 Nice Airport Slide, Cretan Sea, Nankai Trough). However, future work is required to implement the dynamic load adapta-tion routine that modifies the PID settings according to the changing stiffness of the samples.

Acknowledgments The authors want to thank the reviewers Kerstin Lesny and Kate Moran for improving this paper by their helpful comments. This study was funded by the Deutsche Forschungsgemeinschaft as part of the DFG-Research Center MARUM at the University of Bremen.

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