1 LARGE-SCALE PHYSICAL MODEL TESTS ON

LARGE-SCALE PHYSICAL MODEL TESTS ON SAND-FILLED GEOTEXTILE TUBES AND CONTAINERS UNDER WAVE ATTACK P. van Steeg1, E. Vastenburg1, A. Bezuijen1, E. Zengerink2 and J. de Gijt3 Geotextile encapsulated sand elements, such as geotextile tubes or geotextile containers, are considered more and more as an alternative material for coastal protection. Uncertainty with respect to the stability of geotextile elements under wave attack is one of the reasons why these systems are not applied widely yet. Therefore, large-scale physical model tests focusing on the stability of geotextile containers and tubes under wave attack have been performed in the Delta Flume of Deltares. The first part focuses at the stability of geotextile containers under wave attack; the second part describes the stability of geotextile tubes under wave attack. For the geotextile containers, observed failure mechanisms are described and analysed qualitatively. For the geotextile tubes, a physical sound formula based on a theoretical approach and verified with large-scale physical model tests is suggested for both stability and deformation (settlement) under wave attack. It was concluded that Froude scaling is not suitable for geotextile containers (due to sand migration within the containers) but can be applied for tubes with a relatively high filling degree.

INTRODUCTION

Geotextile encapsulated sand elements, such as geotextile tubes, geotextile containers, geotextile bags or geotextile mattresses can be used in several hydraulic applications. Examples and suggestions are given in CUR (2004, 2006), Pilarczyk (2000), and Oh and Shin (2006). However, the elements are not regularly used for coastal defence works. One of the main reasons given in Bezuijen and Vastenburg (2008) is the uncertainty on the behaviour of these elements under wave loads. Therefore, this behaviour is studied by conducting large-scale physical model tests on geotextile tubes and geotextile containers under wave attack. Formulas to predict the stability are scarce. In Pilarczyk (2000) some formulas are given but these are based on small-scale physical model tests. The research described in this paper leads to a formula, based on large-scale physical model tests, providing a dimensionless stability number for geotextile tubes under wave attack and a qualitative description of geotextile containers under wave attack. REVIEW OF STABILITY CONCEPTS OF GEOTEXTILE TUBES AND CONTAINERS

Nine failure mechanisms of geotextile elements under wave attack are discussed in Lawson (2008). Sliding instability, overturning instability, bearing

1

Deltares, P.O. Box 177, 2600 MH Delft, The Netherlands, [email protected], [email protected], [email protected] 2 TenCate Geosynthetics, P.O. Box 236 Almelo, The Netherlands [email protected] 3 Delft University of Technology / Gemeentewerken Rotterdam, P.O. Box 6633, 3000 AP Rotterdam, [email protected]

1

2 instability, global instability, scour of foundation, and foundation settlement are related to external loads. Geotextile rupture, erosion of fill through the geotextile, and deformation of contained fill are internal failure mechanisms. An important, but almost forgotten, failure mechanism is given in Venis (1968). He performed several tests using various sizes of sandbags under current attack. He concluded that, from a certain size, the size at where the sandbag started to shift was independent of the model scale. At a certain critical velocity the sand in the bags started to migrate which led to instability of the sandbags. Venis (1968) concluded that the Froude law scaling is not applicable related to situations where sand migration within the bags occurred. An illustration is given in Fig. 1. Froude law

unstable ucrit Start of sand transport

ucrit,cp

stable L sliding

caterpillar mechanism

Figure 1. Stability as function of element size (represented by L) according to Venis (1968).

Several formulas based on physical small-scale model tests are summarized in Pilarczyk (2000) and CUR (2006). For stability, usually the failure mechanisms sliding instability or overturning instability are considered. The used dimensionless stability numbers are usually based on Froude scaling laws such as Hs/( D) = C or Hs/( B) = C where Hs = significant wave height, B = width of element, D = height of element, = relative density ( sand - water)/ water, = density and C = constant. In Recio and Oumeraci (2009), stability formulas including deformation of the elements are reported. TEST SET-UP AND TEST RESULTS General

Large-scale physical experiments on the stability of geotextile tubes as well as on geotextile containers under wave attack have been performed in the Delta Flume of Deltares. This flume has a width of 5 m, a length of 235 m and a depth of 7 m. Irregular waves, with a JONSWAP spectrum and a wave steepness based on the peak wave period of s0,p = 0.03 were generated at the wave board during each test. The number of waves during each test was approximately N = 1000. In case damage occurred, the test was aborted. If no damage occurred, a new test with a higher wave height was started. During all

3 test series a Geolon® PE180L geotextile was used, thus only woven geotextiles were tested. The opening size of the geotexile pores was O90 = 0.170 mm. Geotextile containers

The width of the containers B was 2.75 m, the length L (perpendicular to the direction of the flume) was 5 m (the same as the flume width) and the average thickness D was 0.55 m. The filling percentage pA, defined as actual fill divided by maximum fill possible with the given circumference of the container, was approximately 44 %. A stack of containers was placed on an existing 1:3 concrete slope. The seaward slope of the containers was 1:2. Above the containers, that formed a berm with a width of approximately 6 m, a smooth 1:3 concrete slope was present. The test program consisted of two subsets, both with a different water level. During the first test series, the water level was 0.75 Hs above the top of the upper container. During the second test series, the water level was equal to the top of the upper container. Measurements that were performed are (i) wave measurements (wave gauges), (ii) profile measurements of the structure (mechanical profiler), (iii) registration of water movement in the container (electromagnetic velocity meter) and (iv) resistance measurements (as measure for the density of the fill material) by pressing a cone through the sand before and after each test (Penetrologger). Table 1. Test program geotextile containers. Test Rc/Hs (-) Hs (m) Tp (s) N (-) Hs t 1-1a 0.75 1.05 4.57 1015 t 1-2 0.75 1.19 5.09 1015 t 2-1 0 0.76 3.97 1014 t 2-2 0 0.92 4.35 973 t 2-2a 0 0.90 4.32 2084 t 2-2b 0 0.90 4.32 2124 t 2-3 0 1.08 4.67 1011 t 2-4 0 1.21 5.05 1036 t 2-5 0 1.34 5.36 1038

(BD) (-) 0.86 0.97 0.62 0.75 0.73 0.73 0.88 0.98 1.09

Hs D (-) 1.91 2.16 1.38 1.67 1.64 1.64 1.96 2.20 2.68

damage minor reshaping damage no reshaping minor reshaping severe reshaping damage damage damage damage

8 6&7

4.25 m

1:2

4 2&3

1:3

1

0.75Hs 1:2

14 12 & 13 11

4.25 m

1 3.15 m

Figure 2. Schematized test set-up of geotextile containers; test set-up prior to test-run; test-run.

The test program is given in Table 1. In the first test series (t1: Rc/Hs = 0.75), two tests were performed. During Test t1-2 damage occurred. In the second test

4 series (t2: Rc/Hs = 0) seven tests were performed. During Test t2-2, little horizontal displacement of some containers was observed. To determine whether this was a time-based process, the test was repeated two times as Test t2-2a and Test t2-2b. It was observed that the processes of reshaping and deformation was indeed time-based. Geotextile tubes

The tested geotextile tubes were placed on a plateau. This plateau was made of concrete, had a horizontal floor located 3.60 m above flume bottom and a seaward slope angle of 1:2.5. Eight different configurations were tested. In five configurations (F1, F3, F4a, F4b and F5), single placed tubes with varying filling percentages and sizes were tested. Configuration T1 consisted of a single tube with a bar placed at the landward side of the tube to simulate a trench or settlement due to the weight of the installed tube. Configuration P2 consisted of two tubes placed behind each other. Configuration P3 consisted of a so-called 2-1 stack; two tubes placed behind each other with a third tube on top. At the landward side of configuration P3, a bar was placed. Just like in practice, the prefabricated tubes were filled hydraulically with a sand-water mixture in-situ. To reduce the friction between installed tube and concrete flume walls, smooth wooden plates were attached to the flume walls. An impression of the filling process is given in Fig. 4 (a demonstration beside the wave flume). The dimensions of the prefabricated (empty) tubes were determined based on the theoretical shape using the Timoshenko method as given in CUR (2006). It turned out that the calculated shape of the tube matched remarkably well with the measured shape of the in-situ filled tube. Measurements that were carried out included wave measurements (wave gauges), profile measurements (mechanical profiler), displacement measurement (using overlay photographs) and sand migration (by colouring the sand in the tube with injections and visually inspect these after testing). Table 2. Overview tested configurations geotextile tubes. D (m)

B (m)

a1 (m)

R100%

pA (%) 66

F1

single tube, low filling rate

0.57

2.19

1.12

0.75

F3

single tube

0.79

2.04

1.11

0.75

80

F4a

single tube, high filling rate, close to seaside

0.82

1.52

0.85

0.57

109

F4b

single tube, high filling rate

0.82

1.52

1.22

0.57

109

F5

single tube

0.74

2.02

0.91

0.76

72

T1

single tube, bar at landward side

0.88

2.03

1.03

0.76

85

P2

two tubes

0.84

1.99

0.86

0.77

77

P3 2-1 stack, bar at landward side 0.70 1.41 0.65 0.57 91 1) a is distance between most seaward point of tube and seaward side of horizontal part of supporting plateau

5

F1

F5

F3

T1

F4a

P2

F4b

P3

Figure 3. Tested configurations for geotextile tubes.

Figure 4. Filling of tube (demonstration besides flume); Wave impact on tube during test; analysis of sand migration (visual inspection after test by opening tube).

All single placed tubes (F1, F3, F4a, F4b, F5 and T1) failed due to sliding instability (in landward direction). For the tubes with a lower filling percentage (e.g. F1), some sand migration, based on the color injections, were observed leading to minor deformation of the tube. For all configurations, some settlement of the tubes was observed. For Configuration P2, failure occurred due to sliding instability of the landward tube in landward direction. For Configuration P3, global instability occurred (a slip circle in seaward direction). ANALYSIS GEOTEXTILE CONTAINERS

During the tests a combination of four failure mechanisms was observed. These are: 1. Erosion of fill through geotextile. 2. Global instability. 3. Sliding instability. 4. Caterpillar Mechanism. The four failure mechanisms are described below:

6 Erosion of fill through geotextile: Based on the sieving curve of the sand and opening sizes of the geotextile, 27 % percent of the sand could theoretically move through the geotextile (assuming no natural filter). Actual erosion of fill, based on comparison of the profiler measurements before and after the test series, is estimated between 0 % and 8 %. Global instability: a slip circle containing containers 2, 3, 4, 6, 7 and 8 (see Fig. 2) was visually observed. Sliding instability: sliding of several containers was observed. Caterpillar Mechanism: The measured deformation was large (based on profiler measurements and visual observation) which is largely due to sand movement within the containers. Penetrologger measurements showed a lower packing density after the tests and velocities of more than 1 m/s were measured at several locations within the containers indicating internal sand movement. It was observed that the deformation led to a new mechanism indicated as Caterpillar Mechanism. This Caterpillar Mechanism is not fully understood but it was concluded that moving sand led to a mechanism that can be compared with the rolling of a caterpillar. It was not possible to quantify the contributions of the four observed mechanisms; the combination of observed physical processes was too complex and the processes influenced each other too much. Based on observations and described processes of Venis (1968), it is very likely that the sand movement inside the containers is not dependent on the size of the containers (assuming not very small experimental set-ups). This implies that the scaling with respect to Caterpillar Mechanism is 1:1. In other words, in a prototype situation, start of displacement would occur at the same wave height as in the large-scale experimental set-up in the Delta Flume. The stability of containers in this test set-up is significantly lower than the stability of containers in earlier smallscale model tests. The difference in stability is explained due to the following two aspects: (i) The migration of sand caused a caterpillar mechanism, which contributed significantly to the instability of the containers; (ii) The presence of a 1:3 smooth slope above the containers causes a severe wave run-down, which might have affected the stability of the containers negatively. ANALYSIS: STABILITY SINGLE PLACED GEOTEXTILE TUBES Theoretical approach

For all configurations, except Configuration P3, horizontal sliding instability was the normative failure mechanism. Usually this mechanism is described using dimensionless stability numbers Hs/( D) or Hs/( B). In Deltares (2010), a theoretical derivation is described which led to the following equation: H rep BD

2

(C D

fCL )

( f cos

sin )

(1)

7 Eq. (1), with representative wave height Hrep, shape factor = B/D, wave velocity coefficient , drag coefficient CD, lift coefficient CL, friction coefficient f and slope angle , forms the physical basis for the chosen dimensionless stability number. Assuming no severe wave run-down the ‘±’ can be replaced by a ‘+’. Since the shape factor is a function of the width B and height D of the tube ( = f(B, D)), the stability number Hrep/( (BD)) is, from a theoretical point of view, not a solid stability number. However, for engineering purposes, this stability number turns out to be very useful as will be demonstrated below. All parameters of Eq. (1) are described in the following sections. Representative wave height Hrep. Breaking waves results into transmitted wave energy Et, and so-called blocked energy Eb, consisting of dissipated and reflected energy. The blocked energy can be splitted into energy components on the tube Eb,tube, and on the structure underneath Eb,struc. Since the interest lies in the wave energy on the tube, the following equation with reduction parameter , representtative wave height Hrep, significant incident wave height Hs, “blocked” wave energy on tube Eb,tube, and incident wave energy Ei is suggested: H rep

Eb,tube

Hs

Ei

(2)

The energy components, with transmission coefficient for a situation with only underneath structure Ctr,struc, and transmission coefficient for a situation with underneath structure and tube Ctr,tube&struc are given by: With tube:

Etr ,tube &struc with Etr ,tube & struc

Ei

Eb,tube

Eb, struc

Without tube: Ei

Eb , struc

Etr , struc

with Etr , struc

Ctr2 ,tube & struc Ei (3)

Ctr2 ,struc Ei

(4)

Combining Eq. (2), Eq. (3) and Eq. (4) gives: Ctr2 , struc Ctr2 , struc

tube

(5)

The transmission coefficient can be determined in various ways. In this analysis, the method described in Van der Meer et. al. (2003) is applied. A graphical representation, which is restricted to a situation with perpendicular wave attack (wave angle = 0) and a water level equal to the level of the highest point of the tube (crest height Rc = 0), of this approach is derived and given in Fig. 5. The representative wave height Hrep of Eq. (1) is described by: H rep

Hs

(6)

8 1.0 p=1 0.8

p=2 p=3

X

0.6

0.4 0.2

0.0 0

0.5

1

1.5

2 H s/D

Figure 5. Design graph for reduction coefficient

2.5

3

3.5

4

(applicable for Rc = 0,

= 0).

Friction coefficient f, slope angle . The friction coefficient f between the tube and its foundation layer can be determined by performing friction tests as described in Deltares (2010). The representative slope angle is known for each situation and can be determined as described in Deltares (2010). Drag coefficient CD, lift coefficient CL, wave velocity coefficient . In case of no lift forces, lift coefficient CL is equal to zero. Eq. (1) is rewritten by: 1 ( f cos CD

H B

2

sin )

(7)

In case of no drag forces, drag coefficient CD is equal to zero. Eq. (1) is rewritten by: H D

1 f

2

CL

( f cos

sin )

(8)

Literature gives some values for drag coefficient CD, lift coefficient CL (e.g. Recio, 2008) and wave-velocity coefficient (according to Pilarczyk (2000): = 1-1.5). Since it is unlikely that drag coefficient CD, lift coefficient CL and wavevelocity coefficient will be determined for each specific situation and the fact that the friction coefficient f has not much influence since the lift forces are relatively small (fCL ), these parameters are collected within ‘dustbin coefficient’ Cx which can be determined based on experiments. Eq. (1) is rewritten by: Only drag

N s ,1

Hs B ( f cos sin )

C1

with C1

Only lift

N s ,2

Hs D( f cos sin )

C2

with C2

Drag and lift N s ,3

Hs BD ( f cos

sin )

C3 with C3

1 CD

(9)

1 f 2CL

(10)

2

2

(CD

fCL )

(11)

9 Test results as function of derived dimensionless stability numbers (Ns,x)

Test results as function of three dimensionless stability numbers are given in Fig. 6. On the horizontal axis, dimensionless stability number Ns,1 (upper graphs), Ns,2 (middle graphs) or Ns,3 (lower graphs) is given. On the vertical axis, displacement per test, made dimensionless with width of tube ( x/B) (left graphs), and cumulative displacement, made dimensionless with width of tube ( xcum/B), (right graphs) are given.

Figure 6. Test results are function of derived dimensionless stability numbers.

In the upper graphs of Fig. 6, it can be seen that there is a data collapse indicating that the chosen dimensionless stability number (based on drag forces FD and no lift forces FL, tube height D has no influence) is a representative stability number. However, Configuration F1 (tube with low filling percentage) seems to be an outlier. Probably this is caused by lift forces, which played a significant role for this specific configuration, where the tube was relatively thin and wide. In the middle graphs, it can be seen that there is almost no data collapse indicating that the used dimensionless stability number (based on lift forces FL and no drag forces FD, tube width B has no influence) cannot be

10 applied. In the lower graphs, data collapse can be seen clearly, indicating that the chosen dimensionless stability number (based on drag forces FD and lift forces FL, width B and height D both included) is a proper dimensionless stability number. The only outlier is Configuration F4a. This can be explained due to the position of the tube at its supporting foundation layer. The tube of Configuration F4a was in a (relatively) more seaward position than the tubes during other configurations. Assuming an accepted cumulative sliding distance of 5% of the tube width (B), the following equation is suggested: Hs BD ( f cos

0.65

sin )

(12)

Eq. (12) is verified by the physical model tests under the following conditions: Crest height relative to water level Breaker parameter Friction coefficient Angle of foundation layer Filling percentage Radius of tube when 100% filled Characteristic distance Accepted relative sliding distance

Rc p,toe

f 0o 66 % 0.57 1.2

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