geosynthetics applications for the mitigation of natural disasters and for environmental protection

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Geosynthetics applications for the mitigation of natural disasters and for environmental protection H. Brandl  Emeritus Professor; Institute for Geotechnical Engineering, University of Technology, 1040 Vienna,  Austria. Telephone: +431 58801 22101, Telefax: +431 58801 22199, E-mail: [email protected] Received 7 July 2011, revised 17 August 2011, accepted 17 August 2011 ABSTRACT: The paper first descr ibes the vers atile applica tion of geosynthetic s for the mitigation of floods, landsl ides, rockfal ls, debris flows and avala nches . It focuse s on dykes or flood protecti on dams , on ge osynthet ic-r ei nf or c ed st abil ising fil ls (up to 130 m height ) and ba rri er dam s. Geosyntheti c-re inforced float ing emba nkments (up to 70 m height) in cree ping slope s and seismic are as sho w cle ar adv ant age s over rigid str uct ure s (e. g. bri dge s) not onl y from a geotec hni cal poi nt of vie w but als o regar ding economy , maint ena nce and envir onment al aspects. Envir onment al  protection is predominantly considered by gaining renewable energy from the ground via ‘energy- geosyntheti cs’. Several other applicati ons are also menti oned. Compaction optimisation and contr ol of geosyntheti c– soil struct ures is recommended by rolle r-int egra ted CCC (cont inuous compacti on control), thus impr oving their beha viour significa ntly. KEYWORDS: Geosynthetics, Natural disaster mitigation, Flood protection, Rockfall protection, Landslide protection, Environmental protection, Floating embankments, Energy-geocomposites REFERENCE: Brandl, H. (2011). Geosynthetics applications for the mitigation of natural disasters and for environmental protection. Geosynthetics International ,  18, No. 6, 340–390. [http://dx.doi.org/ 10.1680/gein.2011.18.6.340] 1. FLOOD PROTECTION 1.1. Gener al Cli mat e change requir es a close coopera tion bet we en hydr ology , hydr o engine ering and geotec hnical engineer- ing. The increasing freque ncy and magni tude of floods indicate that the previous prognoses of such events should  be revised for many regions. Consequently , the improve- ment of old dykes or fl ood pr ot ection dams and the construction of new ones has become essential since the 1990s. Environmentally optimal solutions are achieved by (re-)creating retention areas, combined with dams that can  be overflo wed, thus flattening the flood wave. Besides these permane nt flood protection mea sur es, tempor ary ones are increa singly required for emerg ency situations. Geosynthetic s have also pro ved very suitable for mobile systems. The topic of coastal protection should also be men- tioned in the list of possible applications of geosynthetics in the field of natural disaster mitigation. However, it will not be discussed in this 2010 Giroud Lecture, because it has already been dealt with in the 2006 Giroud Lecture delivered by C. Lawson (Lawson 2008). Quality assessment of existing structures, design aspects of new ones, and the requir ed saf ety factor s depend on  purpose, risk potential and monitoring intensity . Therefore dykes (permanently loaded by water) and flood protection dams (only temporarily loaded by water) should be distin- guished. Most geotechnical aspects, however, are relevant for both. 1.2. Futur e trends and resi dual risks Floods have af fected millions of people worldwide in recent decades. In several regions the magnitude and freque ncy of floo d wa ves have increased dra mat ical ly since long-term measurements and historical reports have existed. Figure 1  shows an example from Austria, where a 2000 to 10 000-year flood event was back- calcula ted from the flood disaster in the year 2002. Such hitherto singular values cannot be taken as design values for flood protec- tion dams, but they underline the need for local overflow crests or spillway sections. Moreover, they clearly demon- strate that a residual risk is inevita bl e despite most costly protective measures. The so called ‘absolute safety’ as frequently demanded by the public, by politicians, by the media or by jurists cannot be achieved in reality. 1.3. Failure modes of dykes and dams Knowledge of possible failure modes is an essential  prerequisite both for a reliable quality assessment of Geosynthetics International , 2011,  18, No. 6 340 1072-6349  # 2011  Thomas Telfor d Ltd 

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Page 1: Geosynthetics applications for the mitigation of natural disasters and for environmental protection

8/10/2019 Geosynthetics applications for the mitigation of natural disasters and for environmental protection

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Geosynthetics applications for the mitigation of natural disasters and for environmental protection

H. Brandl

 Emeritus Professor; Institute for Geotechnical Engineering, University of Technology, 1040 Vienna,

 Austria. Telephone: +431 58801 22101, Telefax: +431 58801 22199, E-mail: [email protected] 

Received 7 July 2011, revised 17 August 2011, accepted 17 August 2011

ABSTRACT: The paper first describes the versatile application of geosynthetics for the mitigation

of floods, landslides, rockfalls, debris flows and avalanches. It focuses on dykes or flood protection

dams, on geosynthetic-reinforced stabilising fills (up to 130 m height) and barrier dams.

Geosynthetic-reinforced floating embankments (up to 70 m height) in creeping slopes and seismic

areas show clear advantages over rigid structures (e.g. bridges) not only from a geotechnical point

of view but also regarding economy, maintenance and environmental aspects. Environmental protection is predominantly considered by gaining renewable energy from the ground via ‘energy-

geosynthetics’. Several other applications are also mentioned. Compaction optimisation and control

of geosynthetic– soil structures is recommended by roller-integrated CCC (continuous compaction

control), thus improving their behaviour significantly.

KEYWORDS: Geosynthetics, Natural disaster mitigation, Flood protection, Rockfall protection,

Landslide protection, Environmental protection, Floating embankments, Energy-geocomposites

REFERENCE: Brandl, H. (2011). Geosynthetics applications for the mitigation of natural disasters and 

for environmental protection.   Geosynthetics International ,  18, No. 6, 340–390. [http://dx.doi.org/

10.1680/gein.2011.18.6.340]

1. FLOOD PROTECTION

1.1. General

Climate change requires a close cooperation between

hydrology, hydro engineering and geotechnical engineer-

ing. The increasing frequency and magnitude of floods

indicate that the previous prognoses of such events should 

 be revised for many regions. Consequently, the improve-

ment of old dykes or flood protection dams and the

construction of new ones has become essential since the

1990s. Environmentally optimal solutions are achieved by

(re-)creating retention areas, combined with dams that can

 be overflowed, thus flattening the flood wave. Besides

these permanent flood protection measures, temporary

ones are increasingly required for emergency situations.

Geosynthetics have also proved very suitable for mobile

systems.

The topic of coastal protection should also be men-

tioned in the list of possible applications of geosynthetics

in the field of natural disaster mitigation. However, it will

not be discussed in this 2010 Giroud Lecture, because it

has already been dealt with in the 2006 Giroud Lecture

delivered by C. Lawson (Lawson 2008).Quality assessment of existing structures, design aspects

of new ones, and the required safety factors depend on

 purpose, risk potential and monitoring intensity. Therefore

dykes (permanently loaded by water) and flood protection

dams (only temporarily loaded by water) should be distin-

guished. Most geotechnical aspects, however, are relevant

for both.

1.2. Future trends and residual risks

Floods have affected millions of people worldwide in

recent decades. In several regions the magnitude and 

frequency of flood waves have increased dramaticallysince long-term measurements and historical reports have

existed. Figure 1  shows an example from Austria, where a

2000 to 10 000-year flood event was back-calculated from

the flood disaster in the year 2002. Such hitherto singular 

values cannot be taken as design values for flood protec-

tion dams, but they underline the need for local overflow

crests or spillway sections. Moreover, they clearly demon-

strate that a residual risk is inevitable – despite most

costly protective measures. The so called ‘absolute safety’

as frequently demanded by the public, by politicians, by

the media or by jurists cannot be achieved in reality.

1.3. Failure modes of dykes and damsKnowledge of possible failure modes is an essential

 prerequisite both for a reliable quality assessment of 

Geosynthetics International , 2011,   18, No. 6

3401072-6349  #   2011   Thomas Telford Ltd 

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existing dykes and dams and for an optimised design of 

new ones, in connection with the application of geosyn-

thetics. Figure 2  summarises the dominating failure modes

for typical ground conditions along rivers (near-surface,

low-permeability sandy to clayey silts underlain by high

 permeability sand or gravel)

•   slope failure due to excessive pore-water pressures,

seepage or inner erosion

•   overtopping of the dyke/dam crest

•   slope failure due to a rapid drop of the flood water 

level

•   hydraulic fracture

•   surface erosion and failure of the water-side slope

due to wave action

•   piping due to animal activities, especially from beavers and rats

•   unsuitable planting of dykes (especially trees with

flat roots).

Hydraulic failure is frequently underestimated, but may

occur in different forms (e.g. Eurocode 7;  CEN 2004).

•   By uplift (buoyancy). The pore-water pressure under 

the low-permeability soil layer exceeds the over-

 burden pressure.

•   By heave. Upward seepage forces act against the

weight of the soil, reducing the vertical effectivestress to zero; soil particles are then lifted away by

the vertical water flow. This ‘boiling’ dominates in

silty-sandy soil, and is combined with internal

erosion (Figures 3a, 3 b and  4).

•   By internal erosion. Soil particles are transported 

within a soil stratum or at the interface of soil strata

(Figures   5   and    6). This may finally result in

regressive erosion, leading to ground failure of the

dyke or dam.

•   By piping. Failure by piping is a particular form of 

internal erosion, where erosion begins at the surface,

and then regresses until a pipe-shaped discharge

tunnel is formed (Figure 7). Failure occurs as soon asthe water-side end of the eroded tunnel reaches the

river bed or bottom of the reservoir. Frequently,

several tunnels develop (Figure 8). This process may

 be induced or significantly promoted by animal

activities, as field observations over many years have

revealed (Figure 2).

Hydraulic failure may reach several tens of metres away

from dykes or dams, as experience has shown. This could  be observed even for low flood protection embankments

with a relatively small hydraulic gradient.

1 10 100 1000 10 000 x -year flood event

Maximum

discharge:m

/s        3 

0

100

200

300

400

500

Flood 2002: ca. 420 m /sBack calculated: 2000 to 10000-year event

3

Hitherto observed event

GEV-LM Extrapolation

Figure 1. Statistics for the annual maximum discharge values

of the river Kamp in Austria (Gutknecht et al.  2002)

Water outflow

Eroded material

Cracks

Low-permeability soil layer 

Low-permeability soil layer 

Low-permeability soil layer 

Low-permeability soil layer 

Low-permeability soil layer 

Low-permeability soil layer 

Low-permeability soil layer 

Low-permeability soil layer 

Sand

Sand

Sand

Sand

Sand

Sand

Sand

Sand

Figure 2. Failure modes of dykes and dams: near-surface

stratum of low permeability (possibly with local ‘windows’).

After VDZ (2002)

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Eurocode 7 (CEN 2004) states that in situations where

the pore-water pressure is hydrostatic (negligible hydrau-lic gradient) it is not necessary to check other than for 

failure by uplift. In the case of danger of material

transport by internal erosion, filter criteria should be

used. If the filter criteria are not satisfied, it should be

verified that the critical hydraulic gradient is well below

the design value of the gradient at which soil particles

 begin to move.

Experience has shown that the magnitude of the critical

hydraulic gradient where internal erosion begins is fre-

quently overestimated. Figure   9   summarises the critical

values on the basis of field observations, geotechnical

measurements, literature and long-term experience for 

different soils. For comparison, the conventional criterion

(icrit ¼  ª9/ªw), Lane’s criterion, and the critical zones after 

Eurocode 7 (CEN 2004) or   Chugaev (1965) respectively

are also plotted in the diagram.

Filter protection is generally provided by the use of 

non-cohesive granular material (natural soil) that fulfils

adequate design criteria for filter materials. Filter geotex-

tiles have been used increasingly since the early 1970s.

Common filter criteria for soils are from Terzaghi and 

Sherard, and for geotextiles from Giroud (2003, 2010) and 

Heibaum   et al.   (2006). All criteria have particular limit-

ations, whereby non-cohesive and cohesive soils have to

 be distinguished. Whereas two criteria are sufficient for granular filters (the permeability criterion and the reten-

tion criterion), four criteria are required for geotextile

(a)

(b)

Figure 3. ‘Boiling’ behind dykes indicates inner erosion and

beginning of hydraulic failure: (a) randomly distributed

multiple boiling; (b) detail of an erosion spot

Figure 4. Boiling ‘volcanos’ after the flood

Phase 1 Phase 1Phase 2 Phase 2

(a) (b) (c)

Figure 5. Hydraulic fracture of dykes or flood protection

dams due to seepage through or beneath the dyke or dam

(Ziems 1967): (a) suffusion (fine particles move into pore

voids of coarse grain fractions); (b) contact erosion at the

interface of soil strata; (c) internal erosion in steady-state

flow condition

Figure 6. Contact erosion perpendicular or parallel to an

interface of soil strata

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filters (Giroud 2010): the porosity criterion and the

thickness criterion also have to be considered.

1.4. Quality assessment of dykes and dams

Comprehensive in situ investigations have disclosed that

the quality of dykes and dams with a height,  H , 15 m is

mostly lower than that of higher earth structures. There-

fore low but long flood protection dams may involve

rather high risk potential. Dams of   H > 15 m, however,

have already been constructed for decades according to

the strict regulations of ICOLD.

When assessing the quality of dykes or flood protection

dams, the purpose and risk potential of the structureshould be taken into consideration. Dykes as fill dams

directly along reservoirs, rivers or canals are continuously

1

6

3

4

1 Free water table2 Piezometric level in permeable subsoil3 Low-permeability soil

4 Permeable subsoil5 Possible well; starting point for pipe6 Possible pipe

2

1

5

Figure 7. Example of conditions that may cause piping (Eurocode 7; CEN 2004)

Figure 8. Piping through a railway embankment during an

already sinking flood

Medium dense – stiff 

     i  c       r

        i        t 

Fine Fine FineMedium Medium MediumCoarse Coarse CoarseClay

Silt Sand GravelCobbles Boulders

1.2

1.1

1.0

0.9

0.8

0.7

0.6

0.5

0.4

0.3

0.2

0.1

0

γG,dst

γG,dst

D

 AE

Hydraulic fracture EN1997-1

i crit

γ

γW

0.001 0.002 0.006 0.02 0.063 0.2 0.63 2 6.3 20 64 100 200

Grain size (mm)

BC

Figure 9. Critical hydraulic gradients for hydraulic fracture (internal erosion) (Brandl and Hofmann 2006);   i crit:  depends not

only on grain size distribution and density/stiffness but also on flow pressure;  ªG,dst   partial safety factor for permanent

unfavourable effects

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monitored, whereas flood protection dams are loaded by

water only during floods. The latter are therefore more

critical, as experience has shown.

Besides levelling of the crest and control of the

geometric data, the following (geotechnical) investigations

have proven successful

  seasonal field observations, especially during floods•   documentation of water discharge and localisation of 

wet slope spots and zones during floods, depending

on the magnitude and duration of the flood wave

•   aerial photographs, especially during and after floods

(and in normal periods for comparison)

•   spot-checking by soundings, borings or exploratory

 pits

•   field tests and laboratory tests

•   geophysical investigations

•   tracer testing

•   infrared photographs for localisation of seepage

•   roller-integrated continuous compaction control

(CCC).

Investigations in steps provide the best results, and are

the most cost-effective. Conventional spot-checking gives

only random data, whereas investigations covering the

entire area provide more detailed information. Conse-

quently, geophysical methods have been increasingly used 

over the last 10 years. However, experience has shown that

a reliable interpretation is possible only if at least two or 

three different methods are applied (e.g. geoelectrics,

micro-gravimetry, combined surface wave/refraction seis-

mic). Geophysical methods and electric tracers are also

used to check the effectiveness of sealing elements (sur-

face sealings, core sealings, cut-off walls). Multi-sensor 

survey systems together with spatially targeted electrical

tracers make it possible to localise leakage areas, and 

hence leaks. Moreover, they have proved suitable for 

continuous leakage monitoring of sealing systems (liners,cut-off walls).

A complete quality assessment of the dyke or dam crest

is made possible by roller-integrated continuous compac-

tion control (CCC). This method registers the interaction

 between roller and ground (for details see Section 5). The

measuring depth reaches 2 to 2.5 m, and thus covers the

most critical zone of dykes and dams. Weak points are

easily localised, which is especially important for struc-

tures with crests that can be overflowed (spillway section).

However, if the crest has a top cover of concrete, asphalt

or riprap, or if it is very uneven, CCC cannot be used.

The advantage of CCC is not only full control but also

a compaction of weak zones. Figure   10   shows some

results along an old dyke that was severely attacked by

the previous flood. The weak points and zones were not

visible along the crest, but could be clearly localised by

CCC during the first roller pass. After six roller passes

the quality of the top 2 m of the dam had improved 

significantly, although relatively weak spots still existed.

However, the control values exceeded the lower limit

value (30 for this particular fill material) with only one

exception (23).

1. Roller pass

6. Roller pass

120

120

100

100

80

80

60

60

40

40

20

20

0

0

110/100

110/100

110/150

110/150

Cursor: 110/190Required limit

Cursor: 110190Required limit

15.930

29.130

Mean value: 38.6SD: 20.5

Mean value: 56.3SD: 18.2

Max: 89.1Min: 10.3

Max: 94.7Min: 23.0

110/250

110/250

110/300

110/300

110/200

110/200

Pt1: HMV 77.8

Pt. 2: HMV 42.7

Pt.3: HMV 29.5 Pt.4: HMV 23.0

Figure 10. Results from roller-integrated continuous compaction control (CCC) of an old flood protection dam after a severe

flood. Weak spots are clearly visible; also improvement after six roller passes. Compaction degree/stiffness (dimensionless value)

is plotted against the chainage of the dam

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1.5. Geosynthetics application for dykes and dams

Most dykes and flood protection dams are lower than 10

to 15 m, but are clearly longer than dams for hydropower 

generation. Moreover, the latter are cross structures

whereas dykes and flood protection dams usually represent

longitudinal barriers. Consequently, different design rules,

maintenance and monitoring aspects have to be consid-ered. This leads to significantly increased application of 

geosynthetics for dykes and flood protection dams.

Geosynthetics have proved successful for emergency

and temporary measures as well as for permanent pur-

 poses, and for contingency plans (successful defence of 

dykes and dams against severe floods). Furthermore,

geosynthetic sensor mats serve for monitoring of critical

zones.

1.5.1.  Emergency or temporary measures

Placing filter stable counterweights (berms) can prevent

hydraulic failure of the dyke or dam by seepage or uplift,

or by internal erosion and piping. Sandbags are preferred 

for local stabilisation, and sheets of filter geotextiles

(covered with sand, gravel, or other granular material) for 

larger critical zones (Figures   11   and   12). Water outflow

must not be prevented, as it would create excessive pore-

water pressures and favour sudden failure. Consequently,

 placing impermeable geomembranes or plugging weak 

spots with clay at the land side is counterproductive

(although it is in many cases the instinctive reaction of 

those defending the dyke or dam).

If the dyke or dam is already broken, progressive failure

can be avoided by using a helicopter to place, first, stiff,

heavy elements (e.g. military anti-tank steel crosses) toform a skeleton, and then fill large bags and finally

smaller sandbags (Figure 13). After the flood, and removal

of the emergency elements, the waste consisting of soil

and geosynthetics should be recycled (Figure 14).

Mobile flood protection systems may be stiff panels

inserted in fixed toe elements and other modular systems

of metal, reinforced concrete or synthetic material.

Figure 11. Hydraulic failure prevention by covering local

water outflow with permeable sandbags (small geotextile

bags)

Figure 12. Hydraulic failure prevention (emergency measure)

by covering the entire dam slope and berm with gravel

placed on a filter geotextile

Figure 13. Large bags and sandbags between anti-tank steel

crosses to close a broken dyke section (emergency measure).

Sequence of placing: 1, steel crosses; 2, large bags; 3.

sandbags

Figure 14. Soil–geosythetics waste after the flood when

removing the emergency elements

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Geosynthetics are a promising alternative whereby semi-

 permeable and impermeable systems have proved success-

ful. Geotextile container solutions show a great versatility:

small bags, large bags, containers, tubes and mattresses.

Many of them are also used for permanent purposes

(Saathoff  et al.  2007).

Pre-filled sandbags, which can be easily carried and 

 placed by one person, are frequently preferred to larger geotextile sand containers. They have already been suc-

cessfully used for decades. However, the better the access

and placement conditions for technical equipment, the

larger the containers that can be used. Finally, if there is

no access to critical zones on land, large bags can be

 placed by helicopter.

In all cases the geotextile should have high robustness

(puncture resistance), high elongation behaviour and inter-

face friction. Flexible behaviour allows the geotextile

container to mould itself in with the existing features, and 

also allows a certain degree of self-healing of the

structure. Unavoidable damage from driftwood and similar 

objects is also then minimised.

The permeability of structures made of geotextile sand 

containers depends primarily on the size of the container 

elements and the method of placement. The grain size

distribution of the infill is negligible.  Recio and Oumeraci

(2008) showed that the flow through the structure is

governed solely by the gaps between neighbouring con-

tainers, and that the flow through the sand fill in the

containers can be neglected. Figure   15   compares two

models (out of 11) providing the worst and best results.

The permeability coefficients obtained are in the order of 

5 3 10 –2 m/s and 7 3 10 –3 m/s, respectively, despite a

rather similar width of the structures. Furthermore, place-ment of the containers such that the contact areas between

containers were maximized resulted in the highest stability

against wave action.

Figure   16   shows a modular system of metallic grid-

 boxes lined with nonwoven geotextiles and filled with

sand. Such elements are preferably used along roads and 

in settlements and cities. By placing them both side by

side and on top of each other, massive and high barrier 

structures can be constructed. Water discharge is largely

negligible, roughly comparable to that through a wall of 

closely placed geotextile sandbags.

Synthetic impermeable tubes (usually of PVC) are easy

to transport. They are connected by straps, then inflated 

with air, moved to a fixed position and filled with water 

(Figure   17). Geotextile tubes to mitigate the wave-

 breaking process in meandering rivers or to fill broken

sections of dams are less used, but they have proved verysuitable as low-crested submerged structures to reduce the

incident wave energy on shorelines.

Mobile flood protection systems require good and 

repeated training of the emergency staff to assemble and 

activate the barriers as quickly and effectively as possible.

Medium

Cross-section Front view Plan view

GSC, one above theother 

Containers above each other,maximal size of gaps

LargeLarge Large

Model

1

Model11

GSC, overlapped

Medium

MediumMedium

Medium

OverlappedOverlapped

Gap

Figure 15. Models to determine the hydraulic permeability of structures made of geotextile sand containers. Significant

influence of container size and method of placement on hydraulic permeability:   k    53 10 –2 m/s (Model 1),   k    73 10 –3 m/s

(Model 11) (Recio and Oumeraci 2008)

Figure 16. Mobile flood protection system with gabions of 

lined nonwoven geotextiles filled with sand. The structure

finally withstood a 100-year flooding event

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1.5.2.   Permanent measures for dykes and flood protec-

tion dams

Figure   18   gives a schematic selection of geosynthetics

application for dykes and flood protection dams. Not

included are geosynthetics for overflow (spillway) sec-

tions, for vertical cut-off walls, for geotextile sand con-

tainers and tubes. The latter are used for erosion control,

scour protection, scour f ill-in, groynes and breakwaters.

Experience gained in coastal engineering can be widely

adapted to rivers. There are, of course, certain differences

 between constructing new dykes and flood protection

dams and repairing or refurbishing existing structures, but

the basic design requirements are largely the same.

Dykes or dams fully reinforced with geosynthetics, as

indicated in Figure  18, are rather exceptional cases, for 

instance along the outer bank of a river, or at a local

narrowing of the cross-sectional flow. Transition zones

from rigid structures (e. g. culverts) to the flexible dam

fill and overflow sections may also exhibit geosynthetic

reinforcement.

Figure   19   shows versatile application of geosyntheticsfor the rehabilitation of a section of the dyke at the river 

Elbe after the great flood in 2002. Slope sealing at the

water side of the dyke was widely used there, because

geomembranes or geosynthetic clay liners can be placed 

very quickly.

River bank and canal slope protection also has an

influence on possible flood damage. Solutions with geo-

synthetics generally have lower construction and lifetime

costs than rigid structures.

To sum up, geosynthetics serve the following purposes

for dykes and flood protection dams

•   horizontal, vertical and inclined filters and separation

elements

•   dyke slope sealing at the water side

•   reinforcement of f ill, crest zone, access road (for 

dyke defence)

•   surface erosion protection

•   vertical cut-off walls

•   protection against voles and beavers.

Inclined clay cores of flood protection dams commonly

Figure 17. Mobile flood protection system with synthetic

tubes

Dyke/damdefence way

Dyke/damdefence way

Drainagegravel

Drainagegravel

Separation geotextileFilter geotextile

Separation geotextileFilter geotextile

Separation geotextileFilter geotextile

Separation geotextileFilter geotextile

Erosion protection mats

Erosion protection mats

Erosion protection mats

High level

Rip-rap

Low level

Rock fillor gravel

Excavationin steps

Gradient

4

5

1

5

3

4

5

2 6

Vegetation soil cover 

Flood

GeomembraneGeosynthetic clay liner 

Geosynthetic reinforcement

Sand

Figure 18. Application of geosynthetics for dykes and flood protection dams. Schematic examples, not to scale. 1, filter

geotextile for drainage fill; 2, geosynthetic-reinforced dyke/dam; 3, filter geosynthetics below rockfill or riprap; 4, geosynthetics

for the access way for dyke/dam defence during floods; 5, geosynthetics for erosion protection; 6, geomembrane or geosynthetic

clay liner

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lead to a lower slope stability than vertical cores or cut-off 

walls. This refers especially to a quick drop of the water 

level, causing superficial sliding of the water-side slope,

leaving the clay core unprotected against erosion. In the

case of geomembranes or geosynthetic clay liners, near-

surface slips are also critical, but can be repaired more

easily than along softened, partially eroded clay.

Hydraulic failure may be prevented mainly by two

measures on the land side of a dyke or flood protection

dam

•   installing trenches or relief columns or drainage

wells

•   filling of berms, thus displacing the possible starting

 point of inner erosion or piping further away from

the structure, and decreasing the hydraulic gradient

at this point. Such berms should be constructed as

access roads for quick and easy dam defence in the

case of severe floods.

In many cases berms merely move the hydraulic

 problem further away from the dyke or dam, and retro-

gressive inner erosion may finally reach it in the long term

(after several floods). Boiling and internal erosion have

 been observed up to 20 to 50 m away from dykes and 

dams, even though they were only 3 to 6 m high. More-

over, wide berms are frequently not possible under 

confined space conditions: therefore drainage trenches are

 preferred in these circumstances. However, trenches ex-cavated in very soft soil collapse immediately before

geotextiles and fill material can be placed. The installation

of trussed retaining panels would be too expensive. These

 problems could be overcome by developing ‘relief granu-

lar columns’, jacketed with a filter geotextile:

Jacketed (coated) stone or gravel columns have been

installed in Austria since 1992. At first they were used 

mainly for drainage purposes, for instance as drainage

walls to improve the stability of old flood protection earth

dams. This method has significant construction advantages

over conventional drainage trenches in loose or soft soil.

In critical cases the coated columns are combined with

other measures for dam refurbishment (Figure   20). Thedrainage material (usually clean 4/32 mm, 8/32 mm or 16/

32 mm grain) is lowered by vibroflotation, whereby the

vibrator is wrapped with a nonwoven geotextile (tied 

together at the toe of the vibrator).

The conventional top-feed process of the vibro tech-

nique is not suitable for jacketed granular columns. In this

case the sophisticated vibroflotation technique with

 bottom-feed vibrators is required. The main advantage of 

this method is that the vibrator remains in the ground 

during installation, making the technique ideal for un-

stable ground and high groundwater levels. The granular 

material is discharged from skips into the chamber at the

top of the vibrator and placed at depth (Figure   21). To

avoid geotextile damage, the vibrator is sometimes first

lowered without the geotextile sleeve into the ground to

displace soil. This has proved suitable in coarse or stiff 

subsoil, but is not necessary in fine-grained soft ground.

A leak in the geotextile is quickly recognised, owing to an

increasing volume of fill material during vibration.

Figure   22   shows an alternative technique for relatively

short gravel columns (up to 6 m). Column excavation is

 performed with continuous-flight auger-piling equipment.

Then a steel tube with a slightly greased surface and 

covered with a filter geotextile is lowered into the ground.

When being withdrawn the pipe is filled with gravel,

whereby the geotextile remains wrapped in the ground,

thus enveloping the gravel.Commonly, mechanically bonded continuous filament

nonwoven geotextiles of polypropylene are used. In the

 Air side

80

75

70

Stone cover ( 60 cm)on filter nonwoven

Sand–gravelwrapped into filter nonwoven

Geogrid

Dyke defence path

Old dyke

Dyke creston filter nonwoven

V  : H 1 : 3 

 

  V :  H  1 :  2 

Water side

Vegetation soi l ( 25 cm)Cover soil ( 75 cm)Geosynthetic clay liner (GCL)Sub-base (old dyke section)

d d 

Design water level

0 5 10 15 20 25 30 35 405

Figure 19. Dyke remediation in a section along the river Elbe (Heerten 2006)

Thin diaphragm wall

Drainage pipeon nonwoven

geotextile

Sleeve reinforcedstone columns(a 1.25 m)

Compaction byvibroflotation

(Silty) sandgravel

Clay, silt

Sand, gravel

Figure 20. Improvement of static and hydraulic stability of 

an old flood protection dam

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case of large construction sites, the coating is already

 prefabricated by the manufacturer and distributed in a

tubular shape with a needle-sewn seam fitting specifically

to the vibrator of the vibroflotation equipment. The

geotextile characteristics listed in Table   1   have proved 

successful.To optimise the design of the position, spacing, depth,

diameter and hydraulic capacity of geotextile jacketed 

relief gravel columns requires numerical modelling for 

three-dimensional unsteady flow conditions. Multi-layered 

soil systems always exhibit permeability coefficients that

are different in the horizontal and vertical direction.

Therefore numerical models and calculations should be

calibrated by site measurements or observations along

 previous projects, or by field testing.

The tops of relief columns should be covered with

coarse drainage material, wrapped in filter geotextiles

(Figures   23   and   24) for longitudinal or transverse drain-

age. This drainage layer should carry an access road for 

easy dam defence in the case of severe floods.

Figure   24   shows a cross-section through a new flood 

 protection dam after removal of the old one, which had 

 been destroyed by a severe flood. The coated gravel

columns (diameter 0.7 m) usually exhibit a spacing be-

Figure 21. Installation of geotextile-jacketed drainage

columns (relief gravel columns) by bottom-feed vibroflotation

Figure 22. Installation of geotextile-jacketed drainage

columns by inserting a sleeved steel pipe into a pre-bored pile

excavation. After pipe withdrawal, the gravel-filled geotextile

hose remains in the ground

Table 1. Geotextile characteristics for jacketed drainage

columns

Property Value

CBR puncture resistance (kN) 3.85

Strip tensile strength (kN/m) 24/24

Elongation at maximum load (%) 80/40

Cone drop test (hole diameter) (mm) 15Mass (g/m2) 325

Thickness at 2 kN/m2 (mm) 2.5

Figure 23. Top of geotextile-jacketed drainage columns

embedded in coarse drainage layer as sub-base of the access

road for quick dam defence in case of severe floods

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tween 1.5 and 7.0 m, depending on local factors (geo-

technical and ecological parameters, infrastructure, risk 

 potential etc.); spacing is commonly about 4 m. The

water-side dam slope is covered by a net for protection

against beavers (see Section 7).

In many regions, the crest levels of dykes and flood 

 protection dams will need to be raised in the future.

However, because of climate change, the prognoses in-

volve several uncertainties. Therefore these structures

should be adapted to allow overflow in spillway sections,

or at least wave overtopping. This minimises random

failures of dykes and dam. The results of comprehensive

field tests on sea defences subject to wave overtopping

(Akkerman  et al.  2007) can also be applied for dykes and 

dams along rivers. Three-dimensional reinforcement grids

with additional soil erosion protection represent a flexible

alternative to riprap or even stiffer structures (Figure 25).

1.5.3.   Prevention of seepage

From strict theory, seepage through dykes or flood protec-

tion dams cannot be prevented, but in practice it can be

limited to nearly zero (considering evaporation etc.). This

can be achieved by

•   ‘homogeneous’ dykes or dams of low-permeability

fill material

•   zoned dykes or dams with clay cores (vertical in the

centre, inclined at the water side)

•   cut-off walls (independent of the fill material).

Geosynthetics have proved suitable as an alternative to

clay cores and diaphragm cut-off walls (slurry trench

walls). Geosynthetic cut-offs are inserted into the dyke or 

dam by vibration, or by lowering them into a deep trench

that is then backfilled (Figure   26). Leak detection is

 possible by checking the hydraulic potential on either 

side of the sealing element, by thermometry, and by

geophysical methods (mainly geoelectrics). Moreover,

innovative systems with integrated leak detectors are

available.

Slurry trench walls or deep-mixing walls, however,

have a higher resistance against beaver attack than

geosynthetic screens; they have a statical function

(although theoretically neglected for thin diaphragmwalls) – for example, where not only a sealing function

is required but also the ability to take higher forces for 

special cut-off walls; and they may reach significantly

deeper into the natural ground to reduce or prevent

seepage beneath the dyke or dam. Furthermore, they have

 proved suitable for defined fracture sections of dams: In

an emergency it may be necessary to open the dam at

specific points in order to avoid irregular failures at

random zones. A domino-like progressive failure extend-

ing from the designed fracture section must be prevented.

Consequently, sheet pile elements are also used for such

structures.

HW100

Low-permeability stratum(with ‘windows’)

High-permeability stratum(aquifer)

Crest way

Thin

diaphragm

wall

Potential line with relief 

Defence way (asphalt way)

Relief trenchesor columns

Figure 24. Standard cross-section of a new 75 km long flood protection dam in Austria

Figure 25. Stiff crest structure of a wave overtopping and

spillway section of a dyke damaged by flood

Figure 26. Backfilled geomembrane cut-off wall for low flood

protection dams

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1.5.4.   Mitigation of flood damage

In Austria nearly 3000 flood protection projects are under 

construction, design and pre-planning until the year 2015.

This requires multidisciplinary cooperation, and also con-

sideration of infrastructure, of environmental protection,

and of local, public and legal aspects. Plans for dam

maintenance, precautionary measures and emergency ac-

tions have been developed since the flood disasters of 2002 and 2006. Contingency plans (based on risk ana-

lyses) comprise scenarios from slight but frequent floods

to worst-case events (maximum credible water levels).

Another essential prerequisite for successful mitigation

of flood effects is the comprehensive education and 

training of task forces comprising authorities, organiza-

tions, professional groups and volunteers. Flood wave

 prognoses for dam overtopping or failure, contour maps of 

water level, and warning, alarm and evacuation plans are

already standard, and are highly appreciated by the

Austrian public.

Each new or refurbished dyke or flood protection dam

will be equipped with a parallel flood defence road that

makes quick access possible, so that necessary mitigating

measures can be taken without delay. Materials for such

measures (mainly filter geotextiles) are stored in the

vicinity.

These concepts correspond to a judgment by the

European Court for Human Rights stating the state’s

responsibility to protect its citizens from natural disasters

to an ‘appropriate’ (not an ‘absolute’) extent.

2. GEOSYNTHETIC-REINFORCED

BARRIERS AGAINST ROCKFALL:MODEL TESTS

2.1. Measures against rockfall

In mountainous regions, large-scale rockfalls have become

an essential element in regional planning (Figure   27).

Well-known disasters have drawn the attention of both the

 public and the authorities to the need for extensive meas-

ures to protect critical areas. Optimised investment

addresses several questions concerning risk analysis, such

as the probability of a severe event and, particularly, the

amount of damage it causes. Taking economical consid-

erations into account leads to a limited risk reduction in

many cases of rockfall protection.

In recent years the prediction of rockfall has undergone

significant improvement regarding the trajectories of the

falling blocks by the use of highly developed computer modelling, which considers such factors as mass, geome-

try, damping and vegetation. Ultimately this yields the

 block’s kinetic energy at any position. In most cases,

though, prediction of the behaviour of an unstable block 

itself involves uncertainties, and therefore often requires a

wide parameter variation in calculation and design.

The development of structures to guard against rockfall

has received an enormous boost over the last 15 years in

the field of netting, where the cost–benefit ratio can be

noticeably improved by the high absorption of deforma-

tion energy. These results were also facilitated by full-

scale fall tests, in which not only were further types of 

netting developed, but also new forms of evaluating the

absorbed forces were applied.

Descoeudres (1997) cites various structures against

rockfall, rated according to their deformation energy

(Figure  28). Among these passive measures against rock-

fall, earthfill barriers and especially reinforced embank-

ment dams (barriers) take up the prime position. In

contrast to flexible netting, the energy-absorbing effect of 

 protective barriers is provided mainly by their mass.

Sufficient space, and a topography that makes the rock-

fall’s ‘transit area’ sufficiently well known, are prerequi-

sites for the choice of an embankment barrier.

Furthermore, the risk and the quantity of the expected rockfall have to justify the costly option of constructing an

earthfill barrier rather than using netting.

As the different trajectories of the falling blocks and 

their rotational behaviour can be predicted to a high

degree by computer modelling, it seems appropriate to

take this into account not only when positioning the

 barrier but also in the design of its upper (mountain-

facing) slope. Obviously, geosynthetic-reinforced slopes

offer more possibilities in ‘shaping’ the embankment

 properly to prevent the blocks from running over the

 barrier. Additionally, the catch basin can be enlarged by

steeply sloped barriers. Questions on the comparability of 

 barriers with regard to their overall resistance against

dynamic impact seemed so far to have remained unan-

swered, and the economical aspects of such structures

have not been addressed.

Full-scale tests commonly require a large effort when

comparing various kinds of earthfill barrier under specific

conditions. Therefore, qualitative and semi-quantitative

model tests have been preferred for parametric studies, and 

to investigate the interacting factors of the geosynthetics

arrangement, anchoring lengths, and degree of compaction.

2.2. Test set-ups and performance

Walz (1982) describes qualitative model tests in soilmechanics primarily as a method to recognize the failure

mechanism of stability problems. Starting from this ‘phi-Figure 27. Severe rockfall along an expressway; cars were

crushed

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losophy’, a series of 20 dynamic 1 g   model tests were

carried out on protective barriers against rockfall, scaled 

at 1:50 (Blovsky 2002). In the soil mechanics laboratory

at the Vienna University of Technology special attention

was directed to the measurement of forces, acceleration

and deformations in order to gain comparable results, and 

to enable systematic parametric studies.

In order to record the dynamic impact exactly, the

rockfall was simulated by a rigid pendulum that contained 

dynamic force and acceleration transducers. For exact

measurements it was necessary to retain the six degrees of freedom that a single falling block has. Interpretation of 

the measured data was supported by deformation gauges

in the embankment, and optical recording by two digital

video cameras. Figure   29   shows the pendulum with its

hemispherical penetration surface, dynamic force and 

acceleration transducers, and a threaded pole for addi-

tional weight.

The models were constructed in a steel frame structure

with formwork boards as side walls for the cross-sections

of the barriers, to serve as parallel guides. To achieve a

uniform density, the model soil (sand with a low degree of 

uniformity) was compacted in thin, laterally boarded 

layers in such a way that the required soil mass could be

weighed and controlled (Figure   30a). Without the form-

work for the embankment the conditions in the impact

area would not have been reproducible.

To achieve a maximum slope inclination of 3:1 in the

Grates

Diagonal ringnets

 Anchored nets

Can

Rock sheds

Earth fill

Reinforced barriers

10 20 50 100 200 500 1000 2000 5000 10 000 50 000

kJ

Figure 28. Structures against rockfall according to their deformation energy (Descoeudres 1997)

 Accelerationtransducer 

Thread polefor additional

weight

Forcetransducer 

Hemispherical penetrationsurface (wooden)

Figure 29. Pendulum simulating the rockfall

(a)

(b)

Figure 30. Construction of the ‘standard’ model for

simulating rockfalls on barrier fill dams

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models, the soil had to be compacted at Proctor water 

content, thus gaining a certain amount of apparent cohe-

sion, which of course would not exist within a non-

reinforced barrier prototype for a long time. These steep

slopes were carried out to compare non-reinforced and 

reinforced barriers.

As already mentioned, the dynamic impact was simu-

lated by a rigid pendulum to provide an exact placementof the transducers. The pendulum was orientated horizon-

tally in half of the barrier height at the moment of impact.

To gain information about the behaviour of the barriers

under that kind of load, each test contained several strokes

(‘rockfalls’), starting with a small impulse and increasing

to a maximum impulse to destroy the model. The loading

history of each test had to be exactly the same, to achieve

full comparability. To control the increasing impulse,

additional weight and/or the release height were varied 

(Figure   30 b). The release of the pendulum from its set

height was effected by a steel wire and a special clamp

that provided release almost without jerking. This method 

was essential, since triggering of the measurement was

executed via the acceleration signal of the pendulum in

order to gain sufficient data from the whole period of 

impact at a maximum measuring rate. Optimising these

 parameters led to a possible maximum measuring time of 

2.5 s at 2400 Hz dynamic rate.

2.3. Parametric studies

The main goal of the qualitative model tests was to

compare the influences of geometry and compaction on

the one hand, and the effectiveness of geosynthetic rein-

forcement on the other hand. Therefore 20 model tests

were arranged as shown in Figure  31   (non-reinforced) and Figure 32 (reinforced). Each group started from so-called 

‘standard’ tests (Nos 01 to 03 non-reinforced, and No. 10

reinforced), in which both slopes were inclined at 4:5, and 

the compaction was 100% of standard Proctor density.

Subsequently, individual parameters such as slope inclina-

tion, degree of compaction and the arrangement of the

reinforcement were varied. Combinations of altered para-

meters were considered in this first series of model tests.

The first three tests were conducted under exactly the

same conditions to verify the reproducibility of the results.

Geometric variations included different inclinations of the

uphill slope from 2:3 to 3:1 (33.78 to 71.68).

The compaction was tested in a first model with 100%

standard Proctor density, in a second with 90% standard 

Proctor density and in a third called ‘compaction in

zones’, with a looser cushion (90% standard Proctor 

density), situated as a ‘buffer’ in front of a well-compacted 

core. All these variations were built up for both reinforced 

and non-reinforced tests. Several alternative arrangements

of the geosynthetic elements were investigated by varying

the anchoring length, the barrier type (sandwich or com-

 pound), and the facing method.

2.4. Analyses and results

The measured data were analysed, first, by a detailed comparison of each registered signal (i.e. force and 

acceleration during impact) for the first three strokes of 

each test, and second by an overall comparison, consider-

ing every stroke until failure of the model.

The detailed comparison made it possible to define a

useful term or characteristic parameter for evaluating the

resistance of earth-fill barriers: Single values, such as the

 peak values of force or acceleration, showed specific

dependences on the condition (local compaction) of the

impact area, especially for the first ‘soft’ impacts of eachtest. Therefore the use of a combined term, such as

impulse or energy, seemed more appropriate in this con-

text. Figure   33   gives an example of the force–time

relation for both reinforced and non-reinforced standard 

 barriers. Whereas the long impact periods for strokes on a

non-reinforced barrier caused low peak values in the force

signal, the reinforced tests showed the opposite behaviour.

Therefore the area beyond the signal (the impulse) could 

 be used to reasonably describe the resistance of the

 barriers.

The evaluation showed that in this example the trans-

mitted impulse of impacts on reinforced barriers attained 

higher values than on non-reinforced ones. This result was

also obtained when comparing other data from reinforce-

ment tests. Analysing compaction tests in detail showed 

that proper compaction of the barrier caused better shear 

resistance and consequently a higher impulse at impact.

The effects of the absorbed impulse can be clearly

visualised by comparing for example a reinforced model

with its non-reinforced equivalent from the upper camera

 position, as shown in Figures 34a and  34 b.

The results of the detailed comparison were similar 

when evaluating not only the first three strokes but also

the number of strokes that was required to destroy the

model completely. First, the number of strokes differed for each of the tested variations, and second the measured and 

calculated data for each impact yielded different values.

Figure   35   summarises the overall sum of impulses for 

each tested model, whereby the value of the non-rein-

forced standard barrier was assumed to be 100%.

Most remarkable was the overall difference between the

reinforced and non-reinforced barriers. Clear relations

with the geometry and the mass could also be obtained:

The steeper the slope inclination, or the lower the mass,

the lower the impulses that were registered.

Lower compaction always caused lower resistance of 

the barrier, but the difference from the standard values

was more significant for the non-reinforced models.

Two variations were clearly ahead of all others, namely

the compound type and the sandwich type. Those models

were designed to take the highest loads in two different

ways: The sandwich type required a lot of geosynthetic

reinforcement to cover each layer completely, whereas the

compound type used a relatively small area at the facing,

where the upper and lower ends of the geosynthetics were

 bonded.

2.5. Economic efficiency

To include economic matters, cost calculations for the

corresponding prototypes of the model barriers werecarried out. The following assumptions simulated a parti-

cular project

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Test 15GTX compound type

Test 17GTX 2:1

Test 12GTX without facing

Test 14GTX sandwich-type

Test 16GTX 1:1

Tests 10–11GTX standard/GGR

Test 13GTX short anchoring

10

5

10

10

10

10

10

 4  :  5

 4  :  5

 4  :  5

    2   :    1

 4  :  5

  4  :   5

  1  :  1

     3    :     1

4   :  5   

4   :  5   

4   :  5   

4   :  5   

4   :  5   

4   :  5   

4   :  5   

50

25

50

50

50

50

50

50

135.0

67.5

135.0

97.5

135.0

10

100%

100%

100%

100%

100%

100%

100%

100%

35

8.75

35

35

35

35

3589.2

135.0

122.5

Test 18GTX 3:1

Test 19GTX compaction low

10

 4  :  5

4   :  5    50

135.0

90%

35

Test 20GTX compaction in zones

10

 4  :  5

4   :  5    50

135.0

100%

35

4   :  5   

    3   :    2

   9   0   %

Figure 32. Cross-sections of the reinforced models of barrier f ill dams; percentage of compaction refers to Proctor standard

density; dimensions of dams in cm

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•   barrier length of 200 m, constructed in 0.5 m thick 

layers

•   fictitious construction time of 7 months

•   0.3 m overlapping of the geosynthetic reinforcement

•   2 km transport distance for the fill material

•   no direct passing over the reinforcement by skip

lorries

•   wedge of humus for a vegetated facing of the

reinforced barrier zone.

For the reinforced structures, obstructions between the

filling and reinforcing working teams on the site were

taken into account: Because of the barrier’s geometry,

filling with constant efficiency yields an increasing head-

way of the filling crew. To maintain sufficient utilisation

of the expensive filling machinery, the reinforcement crew

was accordingly resized. Even so, a certain loss of 

efficiency had to be considered when filling the top

(narrow) layers.

Figure 36 illustrates clear relations between volume and 

cost, as well as an approximately 40% higher cost levelfor reinforced barriers compared with non-reinforced ones.

Filling without high-quality compaction does not yield 

significantly lower cost, because the vibrating roller is

neither an expensive equipment nor on the critical path

regarding efficiency. The remarkably higher level of the

sandwich-type barrier results from the higher amount of 

reinforcement in all fill layers.

To outline the economical efficiency of the barrier 

variations, a value–cost ratio was finally determined by

dividing the sum of the impulse by the calculated cost.

Figure 37  shows the results:

The generally higher impulse/cost level of the rein-

forced structures emphasises that geosynthetic reinforce-ment is more effective than additional expense. Bonding

layers for a compound facing provides a much higher 

Non-reinforced, 1st impact

Non-reinforced, 2nd impact

Non-reinforced, 3rd impact

Reinforced, 1st impact

Reinforced, 2nd impact

Reinforced, 3rd impact Impactforce(N)

Impulse (Ns)

Impact period (s)

900

800

700

600

500

400

300

200

100

0

0 10 20 30 40 50 60 70 80

Impact period (ms)

Impactforce(N

)

Figure 33. Example of force– time relation for the first three impacts (reinforced/non-reinforced)

(a)

(b)

Figure 34. Cracks in (a) a non-reinforced and (b) a

reinforced model caused by the impact of the rockfall

pendulum

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100

128

8063

42

66 71

297 298

261

228

348

437

202

162

143

230

262

Sum

ofimpulseaspercentageof

standardtest(%)

Non-reinforced

450

400

350

300

250

200

100

50

0

(01-03)Standard

(04)Slopeinclination2:3

(05)Slopeinclination1:1

(06)Slopeinclination2:1

(07)Slopeinclination3:1

(08)Compactionlow

(09)Compactioninzones

(10)GTXstandard

(11)GGRstandard

(12)GTXwithoutfacing

(13)GTXshortanchoring

(14)GTXsandwichtype

(15)GTXcompoundtype

(16)GTXslopeinclin.1:1

(17)GTXslopeinclin.2:1

(18)GTXslopeinclin.3:1

(19)GTXcompactionlow

(20)GTXcompactioninzones

Reinforced

150

Figure 35. Sum of impulse for all tested models as a percentage of the non-reinforced standard test

Cost(

millions)

Non-reinforced

16

14

12

10

8

6

4

2

0

(01-03)Standard

(05)Slopeinclination1:1

(06)Slopeinclination2:1

(07)Slopeinclination3:1

(08)Compactionlow

(0

9)Compactioninzones

(10)GTXstandard

(12)GTXwithoutfacing

(13)GTXshortanchoring

(14)GTXsandwichtype

(15)GTXcompoundtype

(16)GTXslopeinclin.1:1

(17)GTXslopeinclin.2:1

(18)GTXslopeinclin.3:1

(19)GTXcompactionlow

(20)G

TXcompactioninzones

Reinforced

7.5

6.8

5.55.1

7.4 7.5

9.48.8 8.7

15.1

9.4

8.7

7.46.9

9.3 9.4

Figure 36. Calculated cost in millions of euros

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resistance against dynamic impact at comparatively low

cost. The effectiveness of sandwich-type barriers is rel-atively small, owing to the very high amount of reinforce-

ment; its value ranges even behind the 1:1 inclined 

variation with a lower mass. The barriers with low

compaction (tests 08 and 09) rank at last but two and last

 but one position. From the detailed comparison and video

analyses it can be concluded that a lack of shear resistance

obviously causes the poor results for these barrier types.

2.6. Practical conclusions from model tests

Twenty qualitative model tests on rockfall protection

 barriers provided detailed information about their defor-

mation and failure behaviour. The effects of reinforce-ment, compaction degree and geometry/mass could be

clearly determined. From the applied measuring equip-

ment, the exact high-dynamic registration of the impact

 process provided most significant results, supported by

single-frame analyses of video recordings.

To evaluate the overall resistance of such barriers, the

sum of the impulse proved to be a reasonable term. The

test series on different structures showed a clear advantage

of reinforced barriers, which could also be verified by

comparative studies considering the economical aspects.

In addition to a wider load distribution, reinforced 

 barriers offer more possibilities for a proper design of the

uphill slope and the protection against overtopping. Thetests emphasised that, predominantly, the layer that was

directly exposed to the impact was stretched. By creating

a compound between the adjacent geosynthetic inclusions

(sewing, welding, bonding, etc.), the upper and lower filllayers could be additionally activated as load distributors.

If no compound type is used, both the upper and lower 

anchoring length should be sufficiently dimensioned,

otherwise heavy dynamic impacts would overstretch the

reinforcement.

Even a simple placement of geosynthetics without cover 

on the slope facing (test 12) improves the resistance of the

structure significantly. If a short construction time is

required, this could be a reasonable alternative or compro-

mise.

Comparing just the reinforced alternatives, it has to be

stated that the filling volume/mass cannot be substituted 

 by a special arrangement of the reinforcement (except for 

the compound types). Nevertheless, reinforced barriers of 

lower mass showed a higher resistance than non-reinforced 

 barriers with a higher mass.

Regarding the orientation of the geosynthetic strips,

static and dynamic effects have to be distinguished. The

 best orientation from a static point of view is transverse to

the barrier, according to reinforced earth structures or 

retaining walls. Overlapping is not then required in the

main direction of tension.

In the case of a rock impact, however, the reinforcement

is strained mainly in the longitudinal direction. A trans-

verse orientation of the geosynthetic strips would therefore be disadvantageous A compromise is an alternate placing

of transverse and longitudinal reinforcement in subsequent

Non-reinforced

0

(01-03)Standa

rd

(05)Slopeinclination1:1

(06)Slopeinclination2:1

(07)Slopeinclination3:1

(08)Compactionlo

w

(09)Compactioninzon

es

(10)GTXstanda

rd

(12)GTXwithoutfacing

(13)GTXshortanchoring

(14)GTXsandwichtype

(15)GTXcompoundtype

(16)GTXslopeinclin.1:1

(17)GTXslopeinclin.2:1

(18)GTXslopeinclin.3:1

(19)GTXcompactionlo

w

(20)GTXcompactioninzon

es

Reinforced

100

Sum

impulse/costaspercentage

ofstandardtest(%)

350

300

250

200

100

50

150

88 85

61 68 71

237

223

196

172

347

174164

154

184

209

Figure 37. Value/cost ratio as a percentage of the non-reinforced standard test

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fill layers. Moreover, an isotropic behaviour of the

geosynthetics would be favourable, and various kinds of 

connection instead of overlapping should be applied.

Finally, the models tests showed that the zoned dams (e.g.

tests 09 and 20) did not perform better than the homo-

geneous fill dams. A kind of loose ‘cushion zone’ (buffer)

would possibly cut the peaks of the internal forces, but

decrease the barrier’s resistance against puncture or break-through. If barriers are designed with an extremely slender 

cross-section to save space, the stability against over-

turning and internal restraining forces of the structure

decrease.

3. GEOSYNTHETIC– SOIL BARRIERSAGAINST ROCKFALL, AVALANCHESAND DEBRIS FLOWS: DESIGN ANDCASE HISTORIES

3.1. Design assumptions

The design of geosynthetic–soil barriers against rockfall,

avalanches and debris flow involves inherently more

uncertainties than for other geotechnical structures, espe-

cially if they are situated in seismic areas (seismic aspects

are not discussed in this paper).

Superimposed on the usual uncertainties regarding

scatter and stress-dependent change of soil parameters and 

interactive/composite effects between soil and geosyn-

thetics, is the imponderability of the dynamic impacts,

magnitude and frequency of disastrous events. For in-

stance, rock blocks weighing several tonnes (Figure   27)

with fall heights of 100 m and more have a fall energy

that cannot be absorbed by any protective structure with-out material plasticization and local damage. Accordingly,

flexible geosynthetic-reinforced barriers or covers have

clear advantages over rigid protective structures. They are

designed after the semi-empirical design method based on

calculated risk and contingency plans (Brandl 1979):

Assuming the worst ground parameters and maximum

credible natural event would make it impossible to achieve

the theoretically required safety factors in many mountai-

nous regions. Consequently, local damage is accepted, but

the possibility to easily repair or strengthen the structure

must be provided. In following this design philosophy, the

safety factors may fall to   F  ¼ 1.0 at the moment of arockfall impact causing large deformations. Rock penetra-

tion of 1 to 2.5 m into a geosynthetic-reinforced barrier 

dam can be repaired more easily than severely fractured 

reinforced concrete barriers.

The so-called ‘acceptable safety factor’ for permanent

loads must be higher, depending on the risk of failure, the

failure potential and other factors (as in Section 4). The

acceptable safety factor, commonly used for embankment

stability as well as in practically all types of structural

instability, is an empirical method to ensure a level of risk 

of failure considered acceptable by the owner, and the

 public. After all, no structure is absolutely safe, and the

objective of a design is to ensure an acceptable level of risk of failure rather than a certain safety factor.

 Numerical modelling of rockfalls requires knowledge

about the penetration depth of rock boulders into the

 barrier structure, and about the dynamic forces resulting

from the impact. Both depend on the fall height, the

 boulder mass, and the indentation resistance of the

structure.

For barrier dams against avalanches, different design

assumptions are relevant than for rockfall protection dams.

They are quite similar to those for barriers against mud-flows or debris flows. Mass, density and front velocity at

impact are the dominating parameters. Values of more

than 20 m/s have been observed in many cases.

Large-scale tests with avalanches in the field are usually

impossible, owing to the high risk potential. Therefore

computer-aided events are simulated, thus providing valu-

able data for design of barrier systems. There are several

types of avalanche, with widely differing dynamic charac-

teristics: they include high-speed dust avalanches (Figure

38) and lower-speed wet snow avalanches of high density.

Avalanches do not create such excessive local energy

 peaks on the barrier facing as huge blocks of rockfall:

there is, instead, a quasi-uniform load distribution. Never-

theless, high puncture resistance of the geosynthetics is

required, because avalanches may include tree trunks,

which can perform like spears.

Common material models in geotechnical engineering

are suitable for describing the behaviour of soils, but not

the complex interaction of various types of soil and 

geosynthetics. Therefore reliable numerical simulations

require both adequate material modelling and identifica-

tion of the material parameters involved. The scattering of 

material parameters in geotechnical engineering renders

the parameter identification process a challenging task 

(Pichler 2003). Because of the non-linear behaviour of theshear parameters of soil reinforced with geosynthetics, it

is recommended that the parameters be determined under 

conditions similar to those in the field (stress level,

compaction degree).

Many soils exhibit a decrease of shear strength with

increasing shear deformation until the residual value (r ;

Figure 38. Dust avalanche near a high concrete dam

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cr  ¼  0) is reached. This may also occur along soil/

geosynthetic interfaces. Impacts from avalanches and 

especially from rockfall are very short. Consequently,

despite potentially large deformations of barrier structures,

the calculation may be based on design shear values of 

r  ,

design ,

 peak    (1)

In most cases a design value close to the peak value is

tolerable, provided that the decrease of the friction angle

and residual shear angle with increasing normal stress is

considered. This refers to soils as well as to geosynthetic

interlayers and interface friction.

Usually, the uphill slope of barrier dams should be

clearly steeper than the downhill one. This provides a

larger catch area, and reduces the risk of crest overrunning

 by rock blocks, snow or debris. Additionally, catch fences

are installed on most barrier dams. This causes strong

stress constraints at the crest, requiring particular geosyn-

thetic reinforcement. The fences themselves are occasion-

ally of synthetic material, but usually of high-tensile steel

wire (coated nets and meshes, etc.). At present, the maxi-

mum retention capacity of fences is 5,000 kJ. Synthetic

fences are used only on top of critical mountain zones to

avoid the formation of avalanches (Figure 39).

Figure  40 shows different arrangements for reinforcing

steep soil dams and embankments. Usually, the geosyn-

thetic reinforcement exhibits equal spacing and length of 

the geosynthetic inclusions (Figure   40a). An irregular 

spacing pattern reflects those cases where stresses are to

 be expected higher on the top than in the lower regions

(Figure  40 b). Short edge strips (secondary reinforcement) provide surface protection if the spacing of the geosyn-

thetic layers is large (Figure   40c). A similar effect as in

Figure 40 b can be achieved by keeping the layers equally

spaced but varying the length (Figure   40d); short facing

layers serve as surface protection, and also facilitate

surface compaction to achieve high compaction at the

edge of the slope.

The focus in the design of geosynthetic-reinforced 

dams, embankments or barrier structures is on internal

and external stability, considering site-specific aspects

such as surface and facing details.

3.2. General stability considerations

There are different possibilities for designing geosyn-

thetic-reinforced barrier dams. In this section a sophisti-

cated method is developed. It was derived from

conventional calculation methods for non-reinforced bar-

rier structures (Brandl and Adam 2000). The design

 progresses in steps, as follows.

•   Internal stability is first addressed to determine the

geosynthetic spacing, length and overlap. Geometry,

surcharge loads, soil parameters such as angle of 

internal friction and cohesion, geosynthetic para-

meters, and interaction parameters such as adhesion

 between soil and geosynthetic are therefore taken

into account.

•   External stability calculations against global slope

failure, sliding and base failure (especially on soft

soil) have to be carried out in the next step.

Furthermore, a transition from internal to external

slope stability has to be considered. The usual

geotechnical engineering approach to slope stability

 problems is to use limit equilibrium concepts

assuming curved or plane failure surfaces, thereby

yielding an equation for the factor of safety

(Figure   41). The problem can be solved using total

stresses or effective stresses. The use of total stress

analysis is recommended for embankments where

water is not involved, or when the soil is not

saturated. Effective stress analyses are preferred for conditions where water and saturated soil are

involved. Equation   2   for total stress analysis and 

Equation   3   for effective stress analysis describe

limit equilibrium for a geosynthetic-reinforced 

dam. Parameters with an overbar represent effective

values, whereas the same expressions without an

overbar are total values. The factor of safety FS

results

FS ¼X

n

i¼1

 N i tan  þ c˜l ið Þ R þ Xm

i¼1

T Gii

Xn

i¼1

W i sin Łið Þ R(2)

Figure 39. Protective fences on top of a mountain to prevent

the formation of avalanches

(a) (b)

(c) (d)

Figure 40. Various geotextile deployment schemes for

stabilising steep soil dams: (a) evenly spaced, same length;

(b) unevenly spaced, same length; (c) evenly spaced, same

length, with short facing layers; (d) evenly spaced, different

lengths, with short facing layers (after Koerner 1998)

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FS ¼

Xn

i¼1

 N i tan  þ c˜l i

 R þXm

i¼1

T Gii

Xn

i¼1

W i sin Łið Þ R

(3)

For saturated fine-grained cohesive soils whose shear 

strength can be estimated from undrained conditions,

the problem can be simplified. Slices need not be

taken, since the soil does not then depend on the

normal force on the shear plane. Figure  42   shows

details of this situation, which results in equation (4).

FS ¼

cLarc R þXm

i¼1

T Gii

W   (4)

•   Finally, the details of to the surface and facing of the

 protective structure have to be considered. When

using geosynthetics with different properties in the

two directions, it is important to recognise how to

 place the geosynthetics in an optimum way. For two-

dimensional cases, the maximum stress is typically

in the direction of the dam face. For three

dimensional cases – that is, for local impacts – 

 placing in the transverse direction can be moreeffective.

3.3. Impact load on geosynthetic-reinforced barrier

dams

The consideration of local dynamic impacts is essential

for protective embankment dams. On the one hand, it is

difficult to predict the area and the magnitude of the

impact force; on the other hand, exact calculation proce-

dures are complicated and costly. In this paper a calcula-

tion method is presented using physical simplifications

and approximations. The procedure is derived from the

simple case of a wedge-shaped dam cross-section, and can be extended to more complicated cross-section shapes.

3.3.1.   Idealisation of fill dam structure

The dam cross-section is idealised as a wedge (Figure  43).

One slice with constant width in the dam axis is consid-

ered. Due to the ‘stocky’ shape of the vertical beam

 bending deformations can be neglected compared with

shear deformations. Accordingly, horizontal load impacts

cause primarily horizontal shear deformations and hor-

izontal shear forces in the dam.

Considering a differential segment   Ad    of the shear 

 beam (Figure 44), the differential shear deflection d    can

 be written as

d  ¼  F 

 A9G d  ¼  ªd    (5)

where   F   is the horizontal shear force;   G   is the shear 

modulus of the dam; and  A9   is the corrected area, derived 

from dividing the area   A   by the geometric correction

coefficient  k.

Integration of Equation  5,  taking into consideration the

 boundary conditions as shown in Figure   43, results in

Equation   7   for the horizontal displacement according to

the horizontal force on the top of the dam.

ð Þ ¼

ð  A

 F 

 A9   ð ÞG d    (6)

ð Þ ¼  F k A

 A A G   ln

 

 A

(7)

The displacements on the top,  0, and on the foundation

 base,  A, are

0ð Þ ¼ 0 ¼  F k A

 A A G   ln

  0

 A

(8a)

 Að Þ ¼  A ¼  0 (8b)

The first derivation of Equation   7   corresponds to the

rotation of shear  ª

( , )η

n 1

Toe of slope(0, 0)

12 3

12

T Gi 

T G1

T G2

Slice i 

W    i     c  o  s  

θ    i   

W i 

  W i

  i  s   i  n

 θ   θi 

(b)

(a)

n

φ

Figure 41. Details of circular arc slope stability analysis for

(c, ) shear strength soils (Koerner 1998)

( , )η

R    1

2

T Gi 

T G1

T G2

η

C.G.

τ  c 

Figure 42. Details of circular arc slope stability analysis for

soil strength represented by undrained conditions (Koerner

1998)

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9   ð Þ ¼  ª ð Þ ¼  F k A

 A A G 

1

  (9)

The described formulas represent the exact static solu-

tion for the vertical shear beam with variable area

representing a ‘slice’ of the wedge-shaped dam according

to a horizontal load on the dam crown, which serves as

the fundamental solution for the dynamic behaviour con-sidered in the following.

3.3.2.   Rayleigh–Ritz approximation method 

The formulation of the dynamic continuum problem

results in a set of differential equations with an infinite

number of degrees of freedom. The basic differential

equations of such distributed parameter systems, with

associated boundary and initial conditions, even in the

actual case of linear elastic solids, but with non-simple

geometry, cannot be solved in an exact manner. In this

case the Rayleigh–Ritz approximation method is used to

overcome these difficulties: The essential boundary condi-

tions are implemented in the approximation, which is not

a solution of the basic differential equations. The basic

idea is to approximate the displacement   (,   t ) of the

shear-deformable dam by one function separable in space

() and time   q(t ), the so-called Ritz approximation

(Ziegler 1998)

,   t ð Þ ¼  q t ð Þ ð Þ   (10)

where   q(t ) is the generalised coordinate of the single

degree of freedom (SDOF) equivalent system of the

continuum. The function  () is properly selected in order 

to meet the requirements of the essential boundary condi-

tions. The function must necessarily comply with thegeometric boundary conditions, and should as far as

 possible also take into account any dynamic boundary

condition. In the actual case the function   () is selected 

from the exact static solution derived in the section above,

which was determined in a sense of best fit for the

dynamic problem.

The function is normalised in such a way that the

deflection on top of the dam is  (0) ¼  1

ð Þ ¼   ln  0

 A

1

ln 

 A

(11)

From Equation   11   the normalised rotation of shear is

derived as

 ð Þ ¼  9   ð Þ ¼   ln  0

 A

11

  (12)

The approximated rotation of shear,   ª(,   t ), is then

defined as

ª ,   t ð Þ ¼  q t ð Þ   ð Þ   (13)

The original system can be rewritten in an equivalent

system of a Lagrange equation of motion of an SDOF

system, taking energy considerations into account. With

the normalisation of the Ritz approximation, the general-

ised coordinate   q(t ) is well illustrated as the measure of 

the translational motion of an equivalent mass  m, and the

kinetic energy becomes

 E kin ¼ 1

2

ð  A

0

r A   ð Þ _2 ,   t ð Þd 

¼ 1

2 m

_q2 t ð Þ

(14)

From Equation   14   the equivalent mass   m can be

determined by using the Ritz separation approximation.

Integration yields

m ¼  r A A

4 A

ln  0

 A

2

3   2 A  2

0   1 þ 2 ln  0

 A

 2 ln  0

 A

2" #( )

(15)

The potential energy is approximated by the strain

Deformation   ηη

 A    A

0   0

F 0η0

 A A0 0/  A A

 A A( ) /  A Ad   η ( )

η ( )

 A A

η A 0

γ ( ) A

η η  A A( ) 0

η η 0 0 ( )

γ ( )0

Figure 43. Idealised shear-deformable dam cross-section

F    A

 A

d

Figure 44. Infinitesimal shear-deformable element

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energy of an equivalent spring coefficient   k , deforming

according to Equations 10  and  13, to give

 E  pot ¼ 1

2

ð  A

0

GA9   ð Þª2  ,   t ð Þd 

¼ 1

2 k q2 t ð Þ

(16)

Integration of Equation   16   results in the effective

stiffness k , given by

k  ¼  GA A

k A

ln  0

 A

1

(17)

The resulting Lagrange equation of motion of the

idealised dam is that of a linear oscillator with natural

frequency ø0

€qq t ð Þ þ ø20 q t ð Þ ¼  0,   ø0 ¼

 ffiffiffiffiffiffiffik 

m

s   (18a, b)

Knowledge of the motion behaviour of the dam can be

used for dynamic analyses, that is, earthquake calcula-

tions. In the following, the basic solution will be used to

design a dam loaded by a local dynamic impact, such as a

severe rockfall penetrating into the dam. The dynamic

incident is idealised by an inelastic impact at a point.

3.3.3.   Idealised inelastic impact 

Impact is a process of sudden exchange between two

colliding bodies within a short time of contact. With

respect to a single impacted body or structure, loading in

such a process acts with high intensity during this short period of time. As a result, the initial velocity distribution

is rapidly changed. Such rapid loading in the contacting

area is a source whereby waves are emitted that propagate

with finite speeds through the dam body, absorbing the

effective energy.

The most critical case is characterised by a horizontal

impact of a rigid body at the dam crown (Figure  45). The

rigid body, that is, a rock with mass   mS   approaches

the dam with velocity   vS:   A plausible assumption for the

velocity distribution must be made that renders deforma-

tion in the subsequent motion over time. By considering

the static deformation of the linear elastic dam under theaction of a dead weight load   F 0   applied at the crown of 

the dam, pointing in the same horizontal direction, a

compatible velocity distribution can be assumed that has

the same linear distribution. Consequently, the generalised 

velocity after impact can be derived directly from the

Rayleigh–Ritz approximation defined in Equation 10  as

_ ,   t ð Þ ¼   _q t ð Þ ð Þ   (19)

In the case of utmost dissipation, it is assumed that the

colliding bodies do not separate immediately after impact.The surface points of contact take on a common compo-

nent of velocity in the direction of the impact at the end 

of the collision process, to give

_9 0,   t ð Þ ¼   _q9   t ð Þ 0ð Þ ¼  v9S   (20a)

or 

_90 ¼   _q9   t ð Þ ¼   v9S   (20b)

Applying the momentum relation on each partial sys-

tem, taking into account the condition of idealised 

inelastic impact (Equation   10), yields the velocity to

common both bodies as

_90 ¼   _q9   t ð Þ ¼   v9S ¼

  mS

mS þ  m  vS   (21)

Taking into account the conservation of energy, the

maximum deflection   u0 ¼  qu of the dam can be calcu-

lated. Conservation of energy requires

 E 9kin þ  E 9 pot ¼   E ukin þ  E u pot   (22)

whereby it is obvious that   E 9 pot ¼  0 at the moment of 

collision and   E ukin ¼  0 at the moment of maximum deflec-

tion, since deformation velocity is zero. Kinetic energy at

the moment of collision is a maximum, and the potential

energy is a maximum at the moment of maximumdeflection

 E 9kin ¼   12

 m_q92 þ  1

2 mSv9

2S   (23)

 E u pot ¼  12

 k qu2 (24)

Combining Equations   23   and   24   with Equation   22

finally yields the maximum deflection of the dam accord-

ing to a horizontal impact of a rigid body as

u0   ¼  q u ¼

  mSvS

 ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffik  mS þ  mð Þ

p   (25)

The maximum deformation function over the totalheight of the dam is given by equations   7   and   10

combined with Equation 25

Original structure Equivalent system

v s   v sms

m*k *

q 0.

Before impact

q t t u

0 u( ) η

Maximum deflectionafter impact

V q s.

m gV G, , ρ   γ

Figure 45. Original cross-section of a barrier fill dam and equivalent system loaded by an idealised inelastic single-body impact

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u ð Þ ¼  q u ð Þ

¼  mSvS ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffik  mS þ  mð Þ

p    ln  0

 A

1

ln 

 A

(26)

3.3.4.  Magnification factor 

Because of the linear elastic behaviour of the considered system, it is possible to introduce a magnification factor.

Thus a reasonable design method is provided, since static

calculations can be carried out in a first step, and in a

second step these results can be multiplied by the

magnification factor     in order to obtain the maximum

forces (stresses) and deformations (strains) from dynamic

impacts.

It is recommended that static calculations be performed 

introducing a horizontal force   F 0  on top of the dam, with

magnitude

 F 0 ¼  mS g    (27)

Applying this expression to Equation   18a   yields the

maximum horizontal static displacement on the top of the

dam as

stat0   ¼ 0  ¼

 mS g k A

 A A G   ln

  0

 A

(28)

The magnification factor     is defined by dividing the

maximum dynamic deformation (Equation   25) by the

static deformation (Equation 28) to give

  ¼ 

u0

stat0

¼   vS

 ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffik  mS þ  mð Þp  A A G 

 g k A

ln  0

 A

1

(29)

3.3.5.   Limit equilibrium design

Following the considerations above, the impact area

should be designed in such a way that the impact load can

 be distributed to a larger dam region, so that maximum

forces and deformations can be reduced in the near field 

of impact. Nevertheless, non-linear material behaviour and 

ultimate stresses and strains should be considered in an

area near the impact zone. Geosynthetic reinforcement

significantly improves the stability against local failure

caused by heavy impacts.

A finite section of the geosynthetic-reinforced dam

according to Figure  46   is considered. The crown of thedam is horizontally loaded with the static force   mS g 

multiplied by the magnification factor   . Equilibrium in

the horizontal direction means that the resulting shear 

force is constant in every horizontal plane of the dam. The

constant driving force  T S is

T S ¼  T   ð Þ ¼   mS g 

¼ 1

k A   ð ÞG 

d  ð Þ

d   ¼ const:

(30)

 Nevertheless, the shear stress due to the horizontal load 

is decreasing with increasing .

In the case of failure, the driving force equals or exceeds the maximum allowable shear force resulting

from shear resistance of the dam in failure surfaces

(assumed to be vertical) and horizontal surfaces, creating

the body shown in Figure   46. The resistance the of soil

 body is composed of two components; the geosynthetics

contribute the third component.

•   Shear resistance in horizontal plane,   T R 1 (), taking

into account the shear parameters   and  c

T R 1  ¼

 A0

2  1 þ

 

0

r g   0ð Þ tan þ A0

0

c   (31)

The horizontal shear resistance increases withincreasing depth and is a reliable shear resistance

component.

 A0

G

G

l 0

F mDYN0 s.g.  

η

d

T ( )

I ( )

 A( )

I  A

 A A

(a)

Z⊥

G

Z⊥

G

ai .tanα

ai 

T 2.1/cosα

T 2.1/cosα

T 1

α

(b)

(c)

Deformation

η η φ ( , ) · ( )u0 0 u

u0 t t q

γ η ( ) d /d

η η u

u( , ) t t q   φ ( )u

·

η

Figure 46. Limit equilibrium design of a geosynthetic-

reinforced protection dam (the failure body is assumed to be

wedge shaped): (a) cross-section; (b) plan view; (c) maximum

deformation curve

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•   Lateral shear resistance in vertical planes,   T R 2 (),

taking into account the shear parameters   and  c

T R 2   ¼

 l  A

 A

 º0r g tan   1

3  3 3

0

 1

20   2 2

0

þ  l  A

 A

c

2  2 2

0 (32)

The lateral shear resistance also increases with

increasing depth, but loses effect, depending on the

opening angle due to the total separation of the

failure body from the remaining dam at large

deformations. Therefore the lateral shear resistance

force should not be considered in a limit equilibrium

design. Nevertheless, it serves as a ‘hidden’ safety

factor covering heterogeneity and local lower shear 

 parameters.

•   Geosynthetic reinforcement provides a significant

increase of resistance against shear failure. It serves

as a load distributor: thus the affected soil bodyabsorbing the impact energy can be assumed to be

significantly larger. Depending on their stress–strain

characteristics, geosynthetic inclusions are mobilised 

in different states of impact loading of the dam.

Linear tensile stress–strain behaviour of the geosyn-

thetics is recommended, whereby the resulting secant

stiffness should be in the range of the elastic

modulus of soil, to take into account the strain

compatibility between soil and geosynthetics. In this

case, soil and geosynthetics would be mobilised 

simultaneously. Furthermore, different stress–strain

 properties and ultimate tensile strength have to beconsidered. From Figure   46 b it is obvious that the

geosynthetic properties should be equal in both

directions in the case of a local impact. If there are

different properties, the design must be based on the

minimum tensile strength. Consequently, the geosyn-

thetic resistance force,   T R G, can be easily calculated 

as

T R G ¼  a i tan Æ  min   Z normal

G   ,   Z  parallelG

  (33)

A limit equilibrium design requires a clear definition of 

the safety factor. In the case of a geosynthetic-reinforced dam it is proposed to use partial factors of safety   ªS and 

ªR :   Ultimate strength values can be adapted to allowable

values for the design, as

T SªS<

  1

ªR   T R 

1   þ2T R 2

þ 2T R 

G

  (34)

In Figure   47   the driving forces   T S due to a dynamic

impact on the dam crest and the resisting forces   T 1R   are

shown in relation to the dam height. In the top area the

safety requirements are not met: thus a geosynthetic

reinforcement must be installed. The design strength of 

the geosynthetics can be determined exactly for each damregion. It is obvious that the geosynthetic reinforcement

must be concentrated in the top region of the dam; both

the length and spacing can be varied, or geosynthetics with

higher tensile strength can be applied.

When a barrier dam is impacted on a locally limited 

area, primary shear forces are produced that may possibly

cause failure. Furthermore, in a three-dimensional consid-

eration the dam can also be affected by a global bending

moment due to a local impact, especially in the upper 

regions of the barrier. In earthfill and rockfill dams, tensile

and flexural stresses cannot be absorbed, so that failure

may occur due to this kind of load. Geosynthetic inclusions

with high ultimate tensile force placed in the zone of the

dam slope lead to a significant increase in the factor of 

safety. Compound – the capacity of a high shear force

transfer – between geosynthetics and soil is essential.

Geogrids, geotextiles, geonets and geocomposites are

suitable for achieving practicable and reliable solutions.

Furthermore, high friction between geosynthetics and soil

is required in order to transfer the tensile forces into thesoil along the embedded geosynthetics. An approximate

calculation can be performed to estimate the additional

 bending moment taken by the geosynthetic layer. There-

fore a finite section of the dam is considered, as shown

in Figure   48.   The bending moment can easily be calcu-

lated as

 M   ð Þ ¼   Z  parallelG   ð Þ   (35)

whereby  () must be estimated; the dam pressure    PDAM

can be approximated by using earth pressure considera-

tions, taking into account the geometry and the overburden

load.

3.4. Design of geosynthetic–soil barrier dams

Geosynthetics reinforcement of a protective dam improves

the resistance of the structure significantly. Slope angles

can be clearly steeper than for non-reinforced barriers.

Furthermore, long-term surface protection is provided, and 

the facing can be designed in various alternatives. An

essential advantage is achieved by the load-distributing

effect of the geosynthetics, so that the danger of local

damage and overall failure decreases significantly. The

impacted energy is transferred to a larger dam area, so

that stress and strain peaks are reduced.Consequently, the design process involves a modifica-

tion to conventional calculation procedures. As shown in

Geosyntheticreinforcementrequired

 A ‘cohesion’ ‘Coloumb’s friction’

0

T RG R/γ

T · F S

SDY N0 Sγ γ·

T T R1

S,

T R

1 R( )/ γ

Figure 47. Limit equilibrium design: driving forces T S

against resisting forces   T R 1

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the previous sections, the design comprises two steps, with

two different kinds of load 

•   continuous dead loads and ‘quasi-static’ life loads

•   local heavy dynamic impact loads.

In Figure  49   an example of a geosynthetic-reinforced 

 protection dam is shown, specially designed to restrain

rockfalls and other impacts. Geosynthetics are applied in

the dam to meet the following requirements

•   increase of the slope angles (1, 2, 3)

•   global stability of the dam (1)

•   distribution of the impacted load (1, 3)

•   stability against local impacts in the top region (2)

•   steep slope stabilisation (3)

•   facing, surface shaping, and surface compaction

assistance (1, 3, 4, 5)

•   heavy flexural reinforcement covering impacts (4)

•   light flexural reinforcement covering impacts (5)

•   reinforcement of weak and/or unstable foundation

soils (6).

The mass distribution of the dam regarding impact

resistance can be improved significantly by installing a

geosynthetic reinforcement: It facilitates the placement of 

relatively more mass in the top zone than in the lower zone of the protective body. Furthermore, the bottom

area of the dam can be reduced to a minimum while the

dam volume is kept constant, serving as an energy

absorber.

If the topography allows various ground plans the most

effective dam shape is that of a convex curvature. Load is

diverted to the abutments by compression in the arch-like

dam body (Figure 50a) and to the dam base, respectively.

In dams with a straight axis, and especially in concave

curved dams, tensile stresses can be caused by horizontal

impacts, which can be taken only by a tensile reinforce-

ment embedded in the outer slope of the dam (Figure

50 b).

The height of such protective structures is more or less

Cross-section

Situation

η

∆P DAM

∆M 

∆M 

∆M ( ) ( ) δ Z ||G

∆P DAM

δ ( )Z 

||G

Z ||G

Figure 48. Limit equilibrium design: flexural stresses from

bending moment covered by geosynthetic reinforcement

embedded in the outer zones of the dam slopes. The

maximum reinforcement tensile force depends on the

ultimate dam pressure  P DAM

Impact

Soft foundation soil6

1

3

2

4

5 6

1

3

2

4

5

Dam stabilisation and impact load distribution

Impact reinforcement

Wall stabilisation, short facing layers, impact load distribution

Slope surface stabilisation and heavy flexural reinforcement

Slope surface stabilisation and light flexural reinforcement

Increase of bearing capacity;reduction of settlements;increase of global stability;prevention of lateral spreading;reduction of basis deformation

Figure 49. Example of a geosynthetic-reinforced protection dam; effect of different geosynthetic barrier or reinforcement layers

Impact

Geosyntheticreinforcement

Damabutment

(a)

(b)

Load transfer by friction Tensile and flexural

reinforcement

Impact

Geosyntheticreinforcement

Figure 50. Different ground plans of geosynthetic-reinforced

fill barriers (protective dams): (a) convex curved dam;

(b) concave curved dam

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unlimited. Depending on the subsoil, conditions heights

up to an order of 100 m (or even more) are possible.

3.5. Case histories

In 1999 a severe rockfall occurred in Tyrol, Austria, that

required the evacuation of about 300 persons because of 

expected further rockfalls of about 300 000 m3 or even

more (1 million m3

). The protective measures comprised two reinforced geosynthetic barrier dams, the larger one

having a height of 25 m and a crest length of 170 m. The

fill material was partly gained by excavating an uphill

catch basin (Figure   51), and had to be placed within two

months, always under the risk of further rockfalls. Down-

slope of the barriers the natural slope steepened, thus

requiring detailed slope stability analyses and future

monitoring. Figure   52   shows the top zone of the barrier 

dam, which exhibited strong geosynthetic reinforcement,

and also the inserting of a catch fence.

The geosynthetic was a continuous-filament, mechani-

cally bonded (needle-punched), nonwoven geotextile of 

 polypropylene, strengthened with high-strength polyester 

yarns, whereby the orientation of its reinforcement was

 bidirectional. The main geosynthetic characteristics are

listed in Table 2.

The compaction of the barrier dams was optimised and 

controlled by roller-integrated continuous compaction con-

trol (CCC; see Section 5). Finally, erosion protection of 

the barrier slopes was achieved by placing a three-

dimensional mat of extruded polypropylene monofila-

ments, strengthened by an incorporated geogrid.

Figure 53  shows a geosynthetic– soil protective embank-

ment dam against large-scale avalanches in the Austrian

mountains. The upper part of the structure exhibits a

modular (segmental) block scheme consisting of geosyn-

thetic loops. This system acts like a composite body

according to the ‘deadman’ principle, whereby friction

along the anchor elements is far less important than in the

case of conventionally reinforced soil structures. Conse-

quently, numerous site measurements and observations

have shown that the internal stability of loop-anchored 

structures is actually higher than that assessed by conven-

tional calculation. Such walls can be idealised as truss-like

structures, whereby the loops are considered truss ele-

ments under tension, and the soil between the loops and 

modular units represents truss elements under compres-sion. Another calculation method is similar to that of 

cofferdams.

The protective dam/structure of Figure   53   was con-

structed in 1981, and has withstood extreme impacts since

then. Long-term monitoring has confirmed its excellent

 behaviour.

The world’s largest protective f ill barrier against ava-

lanches was recently finished (2010). It is a 600 m long

and 27 m high earth dam of 500 000 m3 in Tyrol, Austria,

Figure 51. Rockfall barrier (geosynthetic reinforced earth

dam) with upslope catch basin for 300 000 m3

  2  :  3

H  25m

Single reinforcement

Twin reinforcement

7.0Catch fence

5.0

2.0

0.5

Geocomposites

Detail

8.0

 1. 0 1.0   0       . 5       

Large-scalerockslide

Erosionprotection mat

4  .5   6.0 m

Figure 52. Protective embankment dam against large-scale rockfalls and landslides; soil reinforcement on upper part with

geocomposites; slope cover with humus and geosynthetic erosion mat

Table 2. Geosynthetic characteristics to Figures  51   and   52

Property Value

Tensile strength (kN/m) 10

Elongation at break (%) 13

Tensile strength at 5% (kN/m) 30

Long-term design strength (kN/m) 31.3

(FS creep ¼ 120 years)Thickness (mm) 3.0

Mass (g/m2) 580

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reinforced with geosynthetics (Figure   54): geogrids of 

high-tensile-strength polyester yarns with polymer coating.

Their quality was adapted to the locally prevailing staticalrequirements (Table 3)

Lost formwork of galvanised steel mesh was used to

achieve locally a facing of 2:1, and an erosion protection

grid covered the vegetation soil.

The design of high geosynthetic-reinforced earth struc-

tures should consider that friction angle     and residual

shear angle  r  decrease with increasing normal stress. On

the other hand, under low normal stress levels, higher in

situ friction angles can be observed than determined in

conventional shear tests. The inclined plane test provides

more realistic results for interface shear strengths, espe-

cially if differentiating between sudden sliding and gradual

sliding (Pitanga et al.  2009).

4. STABILITY OF FLOATINGEMBANKMENTS AND BARRIER DAMSIN CREEPING SLOPES

4.1. General

Floating embankments for roads and highways, and 

(large-scale) toe fills as stabilising counterweights to

unstable or creeping slopes, represent an increasing field of geosynthetics application. Embankments along creeping

slopes have become a promising alternative to bridges if 

Excavation:catch basin

for avalanches

  1  :  1.  2  5

0.66

6.40m

8.58m

5.00 m

0.75

1, 2 3

Tensilestraps

1,2 3 4

S1 S2 S3 S4

9.72 m

(a)

III

II

I

Measuringlevels

0.75

 O r i g.  g r o u n d

Slopeto highway

1  :  1  .5   

1  :  1  .4  3  

S5

0.50

Measuring devicesfor tensile forces

Earth Pressure cells

(b)

2.50

2.50 2

.65

mZ1

Z2 Z3 Z4

S1 S2 S3S4 S5

Figure 53. Protective embankment dam against large-scale avalanches and debris flows; soil reinforcement on upper part with

a loop-anchored wall system: (a) cross-section; (b) plan view with measurement details

Figure 54. Barrier dam against avalanches: 500 000 m3:

 L   600 m,  H    27 m

Table 3. Geosynthetic characteristics to Figure   54

Property Value

Tensile strength (kN/m)

longitudinal 58–168

transverse 30

Elongation at break (%)

longitudinal 10.5

transverse 25

Mesh width (mm)

longitudinal 25

transverse 35–30

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they are not too high or if they are situated in the toe zone

of an unstable slope (Figures   55  and   56). The deeper the

creeping layer reaches, and the higher the seismic activity

of an area is, the more advantages gained by the use of 

geosynthetic-reinforced ‘floating’ earth structures. Some-

times a combination of flexible and rigid elements may

 provide the optimal solution.

A non-reinforced embankment on a creeping slopewould – in the long-term – fail as a result of stress

constraints caused by locally differential deformations

within ground and fill (internal safety). Reinforced em-

 bankments, however, more or less undergo block sliding,

and the internal safety remains widely preserved.

Geosynthetic reinforcement of floating embankments or 

other floating fills serves the following objectives

•   minimising the embankment mass due to steep f ill

slopes

•   creating a composite body that acts like a quasi-

‘monolith’, and can move with the creeping slopewithout (relevant) damage

•   providing sufficient overall stability of the

embankment–natural ground system

•   providing high earthquake resistance.

Monitoring of such structures is therefore inevitable. On

the other hand, they allow construction in areas that

commonly have been considered as too risky. Only the

transition zones from creeping to stable slope sections

require increased maintenance.

Polymer optical fibre (POF) sensors based on the

optimal time domain reflectometry (OTDR) technique

make it possible to measure high strains (up to 40%) fully

distributed along the fibre integrated in geosynthetics for 

structures (slopes and retaining structures, dykes, embank-

ments).

4.2. Dominating design parameters

The dominating parameters for designing floating em-

 bankments and barrier dams in creeping slopes are shear 

strength and groundwater conditions.

Figure   57   shows a histogram that clearly indicates the

important effect of weather on the number and magnitude

of landslides in a particular region: Heavy, long-lasting

rainfalls in spring 1975 caused numerous, disastrous slides

in some regions of Austria, as never experienced during

the past 150 years. They were favoured by a preceding

very wet autumn and heavy snowfalls during winter,

which left the ground soaked like a sponge, and rock 

 joints already filled with water before the heavy spring

rains began. Above all, Figure   57   raises the question of 

worst-case design parameters, and underlines the impor-

tance of semi-empirical design with calculated risk, based 

on the observational method and contingency plans. It

would be uneconomic to construct the most expensive

 protective structures by assuming throughout and super-

 posing the most unfavourable parameters. In many moun-

tainous regions this is even technologically impossible.

Future additional measures – even in connection with

remedial works – have proved far less costly than a fully

engineered design based on high theoretical factors of safety.

Risk assessment and stability analyses of slopes should 

always involve determination of the residual shear 

strength. This is especially important for creeping and 

other unstable slopes. Creeping represents a long-term

 process with progressively decreasing shear resistances,

unlike the sudden deformations caused by rockfall, debris

flow or avalanches. This deterioration depends not only on

Typically60° 

Soil exchange

C r e e p i n g  s l o p e 

Figure 55. Floating embankment in creeping slope

Figure 56. Geosynthetic-reinforced floating embankment,

60 m high, in rugged, steep terrain and seismic zone

Numberoflandslides

250

200

150

100

50

01950 1955 1960 1965 1970 1975 1980 1985 1990 1995 2000

Time (years)

Figure 57. Landslides in Lower Austria between 1953 and

1999. Singular weather conditions in 1975 caused excessive

mass movements

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the soil parameters (mainly grain size distribution, grain

shapes, mineral contents and density), but also on the

degree of saturation, and on the level of effective normal

stress (Figure 58).

Consequently, if the normal stress in shear tests is too

small, the measured value of   r    is not the theoretical

minimum. As   r   of soils and geosynthetics mostly de-

creases with increasing normal stress, the overburdenshould be taken into account when assessing the possible

residual shear strength in the field. Deep-seated slide

 planes are more critical than those near to the surface.

An increasing degree of water saturation favours the

tendency towards slickensides and decreasing   r   (Figure

58). Therefore shear or triaxial tests should be performed 

on saturated specimens to obtain the minimum value of  r 

for lower border analyses.

The rate of displacement has an influence on the

residual strength of a range of soil types, fillings and 

weathered decomposed rock with a high proportion of 

fines. In granular soils or rock fills, the effect of rate of 

shearing on the ultimate strength is negligible. In cohesive

material different behaviour may occur when a shear zone,

formed at a residual strength by slow drained shearing, is

then subjected to more rapid rates of displacement.

Creeping soil slopes close to the limit equilibrium

( F  ¼  1), and exhibiting a low residual shear strength, tend 

towards progressive failure with a gradual transition from

creeping to (sudden) slip failure. The risk of tertiary creep

increases with decreasing   r :   Long-term monitoring is

therefore essential for a reliable risk assessment, and to

start stabilising, strengthening or retaining measures in

time. If sufficient data exist, a creeping factor can be

deduced, and future extrapolation is possible (Figure  59).In the case of geosynthetic-reinforced structures in

creeping slopes, or at the toe of unstable slopes, slip

surfaces running through the ground and the structure

have to be investigated. Therefore the interface shear 

strength (peak and residual value) between geosynthetics

and fill material, and the composite shear strength of the

reinforced fills (compound body shear strength), should be

determined. Floating embankments and barrier dams in

creeping slopes may undergo relatively large deforma-

tions. Consequently, residual strength is an essential para-

meter for risk assessment and proper design.

4.3. Creeping pressure on barrier dams in unstable

slopes

The main factors influencing the creep rate of an unstable

slope are

•   slope angle

•   pattern of discontinuities, rock joint fillings

•   shear parameters

•   water pressure

•   external loads or unloading.

In a slope undergoing creep, retaining structures may be

stressed by a lateral pressure   E cr , which significantly

exceeds the theoretical earth pressure at rest,   E 0:   This

creep pressure,   E cr , may also be considered as ‘sliding

 pressure’ or ‘stagnation pressure’ on a retaining structure.

If we assume a (quasi) cohesionless mass and limit

equilibrium    ¼ , the creep pressure becomes a special

case of an increased Rankine earth pressure according to(Figure 60)

 E cr  ¼  m   ð Þª 12

 h2 cos   (36)

where     is the slope angle;   h   is the height; and     is a

fictitious friction angle, including the effects of water 

 pressure (and small cohesion; Brandl 1980).

The multiplication factor   m() also depends on the

stiffness of the retaining structure. Hence geosynthetic soil

structures attract less creeping pressure than concrete

walls. The enveloping values of Figure   61  were obtained 

from numerous in situ measurements. They have already proved very useful for practical design in sliding areas of 

Austria and elsewhere for more than 35 years.

25

20

15

10

5

0

Residualshearangle,

(deg

rees)

      φ        r

0 200 400 600 800

Normal stress, (kN/m )σ n2

Degree of saturation,

(%)S r 

20

40

60

80

95

Figure 58. Residual shear angle r  against effective normal

stress  n; degree of saturation   S r  as parameter. Results of 

direct shear tests with silty-clayey soil

5 10 50 100 500 1000

12 3 5 10 20

5

10

15

20

25

30

35

Slopemovem

ent,

(cm)

      ∆    x

log time, (weeks)t 

Slope 111.8 cmk cr 

Slope 237.4 cmk cr 

(years)

k cr  ∆ ∆ x x 2 1

log /t t 2 1

1

Figure 59. Steady creeping behaviour of two unstable slopes

in weathered schists; definition of creeping factor

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4.4. Stability analyses of geosynthetic-reinforced

earth structures in slopes

Stability analyses of geosynthetic-reinforced earth struc-

tures in slopes are based mainly on slope failure calcula-

tions, taking into account the retaining forces in the

reinforcing layers and inclusions. Frequently, external/global and internal stability are considered entirely sepa-

rately, in a rather formal way: Global failure mechanisms

commonly assume slip surfaces that run completely out-

side the reinforced earth structure, and thus do not cut the

geosynthetic layers or inclusions. The analysis of internal

safety, however, is frequently based on failure lines

running only within the reinforced earth or slope.

In daily design practice, such a formal separation of 

stability analyses frequently results in investigations being

limited to a very narrow scope, namely failure mechan-isms just outside or (predominantly) inside the reinforced 

 part (e.g. a combination of two sliding bodies with plane

slip surfaces). In many cases, an investigation of potential

failure mechanisms of various shapes and positions run-

ning partly outside and partly inside the reinforced zone is

not carried out.

However, it is specifically such a failure mechanism – 

often called ‘mixed mode’ or ‘compound mode’ – that

frequently turns out to be the most probable form of 

failure, thus providing the lowest factor of safety with

regard to slope stability. Ignoring this may lead to a

critical underdesign of the structure, and hence to an

increased risk of failure.

Therefore the length of the reinforcement should be

calculated from earth pressure theory as well as from

slope stability analyses. A proper design has to consider 

all possible slip surfaces to gain the most critical failure

mechanism. This means that cylindrical slip surfaces have

to be investigated as well as plane, logarithmic, or any

form of combined or polygonal failure surfaces, including

 potential slip surfaces in the ground or within the

geosynthetic-reinforced earth structure (e.g. pre-existing

or fossil slip surfaces in the ground, geosynthetic/soil

interfaces, and structural interfaces). An exclusive reliance

on conventional stability analyses, proving internal and external stability without considering mixed forms of 

failure (compound modes), is totally insufficient, as

numerous failure histories have shown.

4.5. Embankments instead of bridges in creeping

slopes

Modern compaction equipment, optimisation and control,

and geosynthetic reinforcement have opened the possibi-

lity of constructing high embankments instead of bridges

for roads, highways and railways (Figure 62).

Experience has shown that generally valid criteria for 

weighing up the advantages of embankments in compari-son with bridges are rather limited. Whether a bridge or 

an embankment is most suitable has to be specifically

decided for each project. The main factors that influence

the pros and cons of an embankment instead of a bridge

are

•   local situation, including existing buildings or 

settlements

•   geomorphology; stability of existing slopes

•   ground properties, including ground and slope water 

conditions

•   depth of slip surface(s)•   seismic activity

•   allowable total and differential settlements of the

Profileof creep velocity

NV

 A

Creepingslope

Dδ1

δ1

M   V   

M BH B

B   V B

E cr 

E cr 

δ

δ

c

G

 A

D

Q Slideplane

ω

φ1

φ1

φ

θ

Q

G

Retainingstructure

h

Figure 60. Theoretical assumptions for calculating the

creeping pressure   E cr  on retaining structures or protective

embankment dams (barrier fills) in a creeping slope.   V B,   H B,

 M B  are external forces on the top of the structure

4

3

2

1

Multiplicationfactor,

()

    m

      φ  

E mcr,h ( ) cosφ γ φ2

for   φ

20 30 40Friction angle, φ (degrees)

h2

2

   R   i  g    i  d

   F   l  e  x   i   b   l  e

h height (depth)of creeping mass

Figure 61. Influence of stiffness of structure on the creeping

pressure   E cr;      slope angle,     (fictitious) friction angle

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embankment crest (with regard to the evenness of 

road pavements or rail tracks)

•   availability of proper fill material

•   material balance of soil excavation or slope cut and 

of fill volume within a certain construction section

•   length and quality of access ways for the transport of fill materials

•   number, diameter, length and location of possible

culverts in the bottom of the embankment or within

the fill

•   statical system of the bridge

•   schedule of construction operation

•   construction costs

•   costs for long-term maintenance

•   local climate; environmental and aesthetic aspects.

In the case of statically very sensitive slope bridges

(e.g. with continuous girder superstructures), the founda-tion requires a high resisting moment (e.g. large-diameter 

sockets or caissons), sometimes with multiple permanent

anchorage. This means that rather rigid (and deep)

footings have to be designed.

Moreover, such buildings must be protected upslope by

a flexible retaining structure, which acts as a first barrier 

(‘primary’ retaining system) against excessive slope pres-

sures (e. g. Figure 63). This may require very long anchors

(up to 100 m or even more) for fixing them in stable

ground. As slope pressures may change with time, long-

term monitoring of sensitive structures in creeping slopes

is essential.

To sum up, bridges along creeping slopes are very

expensive, have a relatively high risk potential, and require

continuous monitoring and comprehensive maintenance.

Moreover, they are more sensitive to earthquakes than

geosynthetic-reinforced embankments.

When comparing bridges and (reinforced) embankments

along creeping slopes, the following aspects should be

considered.

•   Embankments may replace bridges if the fill does

not significantly increase the slope’s creep rate, and 

the expected deformation remains within an allow-

able limit value.•   The construction costs of (geosynthetic-reinforced)

embankments are smaller than those of bridges.

•   Commonly, conventional embankments are superior 

to bridges only in connection with toe-weighting of 

the creeping slope: that is, if the local geomorphol-

ogy allows a toe fill as counterweight. However, in

the case of very deep-reaching creeping strata,

floating earth structures are superior, because bridge

construction and monitoring become uneconomical.

•   In canyon-like rugged terrain, embankments have

 proved suitable if they are intensively teethed withthe original ground. Geosynthetic reinforcement is

recommended then, at least locally (e.g. Figures   56

and  70).

•   Using lightweight products as fill material reduces

the load applied on creeping slopes, but such

materials exhibit low earthquake resistance, because

they tend to liquefaction (and embankment spread-

ing), especially if they have a uniform grain size

distribution. Moreover, the composite effect between

fill and geosynthetic reinforcement decreases with

lower density.

•   Geosynthetic reinforcement makes a reduction of the

fill mass possible, owing to the steeper embankmentslopes. Furthermore, such composite earth structures

have a significantly higher resistance against earth-

Slope bridge (conventional design)

Highway

Deposit

Original surface

Embankment (’green’ design)

Village

Figure 62. High embankment instead of a bridge for a

highway or railway. Scheme illustrating the acoustic and

visual screening of a nearby village through proper

vegetation of the embankment slope. Also indicated is the

possibility of a steeper embankment slope due to geosynthetic

reinforcement and of a deposit upslope of the embankment

Original surface

 Anchoredwall

Slope wash

Beam

25.5 m

70.0m

13.0m

7.0 m17.0m

Mica schistdecomposed,mylonitic zones

 E x t e n s o

 m e t e r s

 1 2 a n c h o

 r s

Figure 63. Typical foundation and permanent anchoring of a

bridge pier in a steeply inclined, creeping slope. Statically

sensitive superstructure (2.6 km long bridge) completely

separated from the slide-protection measures (tied-back retaining structures) around each bridge pier

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quake, spreading and local failure. The overall slope

stability is higher than for conventional fills, but

nevertheless the creep rate of an unstable slope will

increase unless drainage measures are successful.

•   Dowelling of the creeping zone with large-diameter 

 piles or reinforced concrete sockets beneath the

embankment is another alternative to slope bridges

(Figure 64).•   Running along creeping slopes or crossing them

requires contingency plans for bridges as well as for 

embankments. The structures should exhibit the

 possibility of future strengthening or relevelling if 

the results of long-term monitoring require it.

•   The maintenance requirements of embankments is

clearly lower than for bridges.

•   Embankments are environmentally more friendly

than bridges, and their slopes can be planted. In

many cases the earth structure finally looks like

natural terrain (e.g. Figures 65a, 65 b and  66).

4.6. Long-term performance of high embankments

or barrier dams in creeping slopes

More than 35 years of personal experience with numerous

embankments and barrier dams of 30 to 135 m height and 

a length of about 150 to 700 m can be summarised as

follows.

Usually, the stability factors against slope failure or 

ground failure of high embankments are at their lowest

 just at the end of construction. Excessive pore water 

 pressures may occur in fine-grained, wet-fill zones or in

the natural ground. In the long term a gradual decrease in

safety is possible if 

Highway

Granular surface drain

2  :  3  

Soil exchange

Drainagetrench

Potentialslide surfaces

Nonwoven

geotextiles

Sliding loam

 E v.  a n c h o

 r s

Clayeyslickensides

φr  5°

Depthto45m

Road

Retaining

panels5.5m

6.0 m

6.4 mRip rap(cemented)

Sockets (caissons)6.0 4.5 m

Shotcrete20 cmd 

     46m

O r i g i n a l  g r o u n d 

Figure 64. Highway in an unstable slope with slope dowelling and geosynthetic-reinforced embankment. Cross-section through a

socket wall that exhibits reinforced concrete panels on top of the sockets (caissons) as contingency plan for the possibility of 

later tying back with prestressed anchors

(a)

(b)

Figure 65. Highway embankment, 100 m high, instead of a

slope bridge (with comprehensive retaining measures) on toe

of an unstable slope: (a) in 1980 (immediately after

construction); (b) in 2010

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•   the drainage systems fail

•   low-quality fill material was used, and insufficiently

compacted 

•   the fill material or the natural ground tend to

 progressive failure, owing to a very low residual

shear strength.

Slope stability assessment of high embankments should 

consider a possible decrease of the friction angle in the

lower fill zone, depending on the grain size distribution,

grain shape and strength of the fill material: High over-

 burden causes grain crushing, thus leading to a flattening

of the Mohr–Coulomb rupture line in the     –    diagram.Interface friction between geosynthetics and soil usually

also drops with   n:

High embankments should be designed with benches at

about 20 to 50 m of vertical spacing, depending on the fill

material, slope angle, slope stability and local climate.

These benches should have a width of at least 3 m. They

serve for maintenance, facilitate local strengthening (if 

necessary), and act as a braking zone in the case of local

slope sliding or soil liquefaction (e.g. during heavy rain-

falls or earthquakes).

Benches as well as the toe zone of slopes exhibit local

stress concentration (Figure   67). Therefore these partsshould be reinforced with geosynthetics (Figure   68). In

seismic zones the installation of geosynthetic-reinforced 

zones parallel to the sloped surface of high embankments

is recommended, with width at least 10 m. Barrier dams

frequently should be reinforced over the entire area.

Moreover, it can be useful to install reinforced inter-

layers covering the entire ground plan of an embankment,

for example for slender embankments along steep slopes

and without a toe buttress towards the opposite site of a

valley, or without the possibility of a counterweight fill in

a valley bottom.

Such reinforced interlayers should be at least 3 m thick 

and be placed at a vertical spacing of about a quarter to athird of the maximum embankment height   H , and on top

of the embankment. They serve as load-distributing

members (reducing differential settlements), and increase

slope stability.

In the case of proper fill material and careful compac-

tion, the self-settlements ( ss) of high embankments (i.e.

without ground settlements) typically lie below 1% of the

embankment height   H . If high-quality fill material is

available and intensively compacted, self-settlements may

even be reduced to about   ss ¼  0.1–0.2% of   H . On the

other hand, poor compaction and too many silty-clayey

interlayers (e.g. placed in a sandwich-like mode) may lead 

to self-settlements of about   ss ¼  1–2% of   H   or more,

depending on the proportion, thickness and quality of 

‘soft’ interlayers.

Spreading of high embankments is negligible if they are

 properly interlocked or toothed with the natural ground 

(by cutting steps).Long-term experience with numerous highway embank-

ments has shown that such earth structures often exhibit

Figure 66. Embankment, 120 m high (large-scale

counterweight), instead of a highway bridge in a creeping

slope, 30 years after construction

Bench or slope toe

Figure 67. Stress concentration at the bench or toe zone of aslope: results from photoelastic model tests (Hoek and

Gralewska 1969)

Geosynthetic reinforcementembankment or barrier dam

Erosionprotectionmats

q

Bench zonereinforcement meetsstress constraints

Figure 68. Recommended geosynthetic reinforcement of 

bench zones of high embankments or barrier dams in

unstable slopes; schematic

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significant advantages over bridges. They are environmen-

tally friendly, they facilitate a balance of cut and fill

volume during construction, and they require only negli-

gible long-term maintenance. If proper fill material is used 

and carefully compacted, the differential settlements of the

embankment crest, and hence also of the road pavement,

are usually smaller than for low embankments on soft or 

heterogeneous ground.Until about the early 1990s, continuous compaction

control (CCC) and roller-integrated compaction optimisa-

tion were still in an early development phase, and not

generally available. Nowadays, improved compaction

equipment, technology and control methods clearly facil-

itate the construction of high embankments for highways

(and even for high-speed railways). This could be a cost-

effective, environmentally friendly alternative to bridges

that, at the same time, also provides increased factors of 

safety.

High embankments crossing unstable slopes in seismic

areas should have local reinforcement with geosynthetics.

Such inclusions have proved very suitable, as they sig-

nificantly improve the internal and external (overall)

stability of the embankments. Moreover, they reduce

differential settlements of the embankment crest.

4.7. Case histories

Figure   69   shows the peripheral zone of a 120 m high

embankment that was constructed instead of a slope

 bridge along an unstable slope. The low slope stability

required local reinforcement with geosynthetics; also,

nonwoven geotextiles were placed in the base to provide

sufficient long-term function of the drainage blanket. This

earth structure was constructed between 1978 and 1979;now it can hardly be distinguished from the natural slope

(see Figure   66). Maintenance is minimal (unlike to the

adjacent slope bridges), and the stability of the natural

slope could be increased significantly.

Floating embankments instead of bridges have proved 

suitable also in rugged, steep terrain. Figure  56   gives a

 partial view of a 60 m high reinforced embankment in a

seismic zone. The curved crest connects twin tunnels on

either side of the f ill, which is intensively toothed with

the natural ground. Earthquake-induced deformations

would therefore be of a more minor effect than in the

case of a bridge. Figure   70   illustrates the situation at the

end of construction. To gain sufficient internal and 

external stability, heavy metallic reinforcement (tensile

strength ¼  400 kN/m) was also placed locally.

Geosynthetic reinforcement of floating embankments or  barrier dams becomes especially important in areas of 

high seismic activity, and if the shear strength of a

creeping slope is already progressively decreasing towards

the residual value. This is demonstrated by a spectacular 

case history.

A 100 m high slope cut for a new highway triggered a

landslide, superimposing old creeping. The moving area

finally had a length of about 600 m, a width of 270 m, a

height difference of 330 m, and a total mass volume of 

13 3 106 m3:   Inclinometers disclosed the dominant slip

surfaces at a depth of 83 m, and displacements of 20 mm

 per month there. Furthermore, multiple slip surfaces had 

to be considered. Geological observations identified fis-

H

120m

    

Originalground

Weatheredschists

Possible slip surfaces

Benched

Decomposedwet colluvium

Rockfill(riprap)

Reinforcement

30 m Highway

2  :  3  

1  :  2  

RockfillExtremely heterogeneous fillmaterial (sandy silt to rockfill)placed in sandwich form (not mixed)

Figure 69. Side zone of a 120 m high embankment instead of a highway bridge. Earth structure serves simultaneously as

counterweight for the creeping slope and is intensively toothed with the natural ground

Figure 70. Geosynthetic-reinforced floating embankment,

60 m high, for a highway interchange in rugged, steep

terrain. Gabion facing in the upper zone using sandstone

from the adjacent tunnel excavation. Geosynthetics also used

to stabilise the cuts in the natural slope (see background)

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sures and open cracks up to 2 m wide in the order of 

300 m upslope of the crest of the slope cut.

The entire mountainside where this ‘bit cut’ was

excavated comprises a disordered, ‘chaotic’ mass in which

very weak ophiolites are erratically distributed. Moreover,

the ophiolites are serpentinised, sheared, weathered, and 

with passages transformed to clayey material. A certain

 bedding seems to be dipping into the slope at a gentleangle of approximately 108, but it is intensively tectonised,

and severely cut by step-like discontinuities (Figure   71).

 Numerous discontinuities, more or less parallel to the

slope or cut surface, respectively, govern the local and 

overall stability. Moreover, clayey interlayers (mylonites)

with low residual shear strength contributed to the in-

stability of the ‘big cut’.

Several alternative alignments were discussed for safe

crossing of the geologically sensitive and seismically

active area of this ‘big cut’. Owing to the uncertainties of 

a deep-seated landslide the existing alignment and tunnel

options were excluded.

Given the magnitude of the instability and earthquake

danger, the proposed solution in this area was to shift both

carriageways away from the toe of the ‘big cut’ as far as

 possible, taking into account the highway geometrical

restrictions deriving from the presence of existing tunnels

constructed on both sides of the landslide. This required 

comprehensive measures to reach acceptable safety fac-

tors, whereby geosynthetics were used for reinforcement,

drainage/filters and erosion protection (Figure 71)

•   local slope dowelling (shear keys) with large-

diameter bored piles, barrettes or elliptical sockets

on the toe of the ‘big cut’

•   backfill of the toe zone of the cut, and an additional

counterweight fill of 500000 m3, with each layer 

over the ground view fully reinforced with geosyn-

thetics, to restrict mainly superficial slides but also

deep-seated slides•   partial backfill of the slope cut uphill of the

counterweight to avoid local slides (the toe zone near 

the berm should be fully geosynthetic-reinforced)

•   a counterweight embankment (the ‘big embankment’

in Figure   71) of about 3 3 106 m3, 700 m long and 

maximum 135 m high to increase the overall stability

of the creep area, to prevent deep-reaching slips, and 

to carry the highway after realignment

•   an arch-shaped geosynthetic-reinforced support dam

on the toe of the ‘big embankment’ where the slope

forms a canyon-like valley

•   comprehensive drainage measures (deep-drainage

 borings, drainage trenches, drainage blankets), using

filter geotextiles

•   open cracks treatment, slope reshaping

•   contingency plans to respond immediately to critical

monitoring results.

The counterweight fill, upslope backfill and downslope

 big embankment were placed on granular drainage blan-

kets, using filter geotextiles if locally necessary. Unlike

Ophiolitesserpentinised, sheared,weathered, clayey(bedding not relevant)

Drainageboreholes(filtered)

Backfill

Shear keystructure(old pile rowsintegrated)

Counterweight fill500000 m(reinforced with geosynthetics)

3

‘Big embankment’3 10 m

135 m

6 3

maxH 

Old highway axis

New highway axis

C.Co

C2

n

C2C2µε

n

n

n

nC2

C2C2

n

nC1

n

C2C2?

n

?

n

O.M

Se

SeSe

SeSe

Se

SeSe

?

?

O.M

O.MO.M

O.M

Fk?

n

C2

TE

TE

TE

Figure 71. Geotechnical cross-section through a landslide triggered by a 100 m high slope cut (‘big cut’) and stabilising

measures: counterweight fill and upslope backfill fully reinforced with geosynthetics, ‘big embankment’ includes only local

geosynthetics

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the upslope fills and the downslope arch dam supporting

the ‘big embankment’, the latter was reinforced only

locally, in the bench zones, in sandwich fashion at

vertical-spacings of about a quarter of the maximum

embankment height   H , and within the top 5 m of the

embankment. The upper reinforcement served mainly for 

stress distribution and differential settlement reduction.

The structural reinforcement consisted of flexible, high-strength geogrids manufactured from high-tenacity polye-

ster yarns with low creep and an environmentally inert

coating. The longitudinal tensile strength of the upper 

reinforcement and the structural reinforcement is 45 and 

110 kN/m respectively, and the elongation at break is

12.5%. Additionally, three-dimensional reinforcement

grids with erosion protection were placed on the slope

surface.

Roller-integrated continuous compaction optimisation

and control (CCC) was recommended for all earthworks.

5. OPTIMISED COMPACTION OFGEOSYNTHETIC–SOIL SYSTEMS

The interaction between geosynthetics and soil (or other 

granular fill material) depends mainly on the interface

shear strength, the stiffness ratio, the geometry of inclu-

sions, and the compaction of the fill layers. A high and 

uniform degree of compaction favours composite behav-

iour, thus improving the bearing-deformation character-

istics of the entire system. In addition, homogeneous

compaction avoids local stress concentration and stress

constraints in multi-layered composite systems consisting

of geosynthetics and soil or other granular material.

A high and uniform compaction is especially importantfor reinforced soil barrier dams, floating embankments,

counterweight earth structures and other retaining struc-

tures, but new dykes and dams or their enlargement should 

also be compacted intensively and uniformly.

So far, compaction control has been carried out mainly

 by means of spot test methods (on selected points) to

check the density or stiffness of the compacted layer.

Conventional tests to determine density, and load plate

tests for checking soil stiffness, are based on spot check-

ing by more or less random selection, and they are only

superficial. The uniformity cannot be approved, and weak 

 points are unlikely to be detected by such spot tests. Thisinvolves an unavoidable residual risk. Moreover, all spot-

test methods are relatively expensive and time consuming.

Finally, conventional testing frequently delays construction

work, because construction activities must not be carried 

out in the vicinity of a spot test, as ground vibrations

might affect the test results.

Therefore the conventional methods of compaction

control are no longer sufficient for high-quality projects,

especially for high embankments along steep or unstable

slopes and in seismic areas, for barrier fills, and for dams.

Increasing demands on engineered earth structures require

as much continuous compaction optimisation and control

as possible during the compaction procedure. The roller-integrated CCC-technique, which was first used for road 

structures (Brandl and Adam 1997), represents a distinc-

tive improvement, because all control data are available

during the compaction process and over the roller-com-

 pacted area. Moreover, CCC allows a reduction of the

number of conventional acceptance tests, and provides all

data during compaction. Thus the compaction procedure

can be optimised and expedited.

Composite structures of geosynthetic-reinforced soil

require intensive and homogeneous compaction, wherebythe subgrade should be compacted properly (as far as

 possible). Theoretical investigations, field experiments

and site measurements have shown that there is an

intensive interaction between the soil or other granular 

material (fill layers), the geosynthetic inclusions, and the

compaction equipment. Measuring this interaction pro-

vides an excellent tool for three important goals of 

compaction.

•   Compaction optimisation. This refers to the quality

of compaction, to the required compaction energy

and time, and to the required geotechnical para-

meters of the compacted material. Over-compaction

and re-loosening of layers should be avoided just as

much as heterogeneous degrees of compaction.

Therefore a main goal of the cooperation between

geotechnical and mechanical engineering has been

the development of ‘intelligent’ compaction equip-

ment that reacts to locally varying soil or granular 

material properties by automatically changing its

relevant machine parameters. Rollers with automati-

cally regulating compaction systems (vibratory or 

oscillatory) are a significant step in this direction,

which raises compaction from a mere routine craft to

a scientifically based high-technology process.•   Compaction documentation. The CCC technology

involves an automatic registration of all data in such

a way that the results cannot be manipulated. This

data collection is essential not only for site

acceptance but also for quality control and long-term

risk assessment. Furthermore, such information is

very helpful for future rehabilitation of dams and 

embankments (after floods, earthquakes, etc.).

•   Compaction control. This should be widely per-

formed during the compaction procedure. A calibra-

tion of the control data based on the reaction

 between ground and compaction equipment is

essential. Control tests after compaction should be

increasingly reduced to conventional spot checking,

whereas continuous compaction control (compaction

equipment-integrated) should be promoted.

Soil properties inherently vary within a fairly wide

range. Furthermore, locally different fill materials and 

compaction degrees on the construction site, together with

test uncertainties, increase the scatter of compacted fill

material characteristics. Consequently, reliable quality

assessment and stability analyses may become quite

difficult. Such a situation can be significantly improved 

 by applying roller-integrated continuous compaction con-trol (CCC), which facilitates compaction optimisation

during the compaction procedure:

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Vibratory roller compaction takes place by means of a

vibrating drum, which is excited by a rotating mass.

Oscillation of the roller drum changes, depending on the

soil response. This fact is used by CCC in order to

determine the stiffness of the ground. Accordingly, the

drum of the vibratory roller is used as a measuring tool

(Figure   72). Its motion behaviour is recorded by the

sensor; analysed in the processor unit, where a dynamiccompaction value is calculated; and visualised on a dial or 

on a display unit, where data can also be stored. Further-

more, an auxiliary sensor is necessary to determine the

location of the roller. By means of GPS (global position-

ing system) the position of the roller can be located up to

an accuracy of 5 cm.

The dynamic compaction values have to be calibrated 

on the basis of conventional tests: for example, compac-

tion degree   DPr , density   r, or deformation modulus   E v( E v1   rather than   E v2   values from static load plate tests or 

 E vd   from dynamic load plate tests). The main advantages

of this control method are

•   continuous control of the entire area

•   results available during the compaction process, with

no hindrance or delay of the construction work 

•   optimisation of the compaction work, including

 prevention of local overcompaction (which causes

near-surface re-loosening of the layer)

•   full and permanent documentation of the entire area.

CCC possesses the essential advantage that the measur-

ing equipment can be easily mounted on vibratory or 

oscillatory rollers (smooth rollers or sheepsfoot rollers).Experience has shown that the roller operators and site

supervisors have very quickly familiarised themselves with

this control method. Low-quality rollers, which provide

only low compaction quality, can be eliminated, and the

documented data cannot be manipulated. CCC has proved 

suitable on many construction sites, and has therefore

 been obligatory for several years already in Austria,

mainly for roads and railways, embankments, dykes and 

dams, barrier fills, geosynthetic–soil structures, and clay

liners of waste deposits.

Furthermore, the measuring depth of roller-integrated 

CCC is significantly larger than that of conventional

methods: Whereas density measurements commonly reach

a depth of only 0.1 to 0.3 m, and standard load plate tests

about 0.5 to 0.6 m, CCC reaches to a depth of about 2 to

2.5 m.

Figure   73   illustrates this with a case history: The raft

foundation of a power plant required deep soil improve-

ment and partial soil exchange to minimise differentialsettlements. Soil improvement of the loose river sediments

(sandy silts) was performed with stone columns, which

were then covered by a geotextile and layers of sandy

gravel. The exact location and diameter of the stone

columns was registered during the entire earthwork, and 

could still be observed by CCC on the top layer, which

covered the stone columns by 1.5 m (Figure 73).

Consequently, CCC could easily detect those areas

where geosynthetic reinforcement had locally been

omitted in the sublayer(s) despite the design requirements.

High-quality compaction provides numerous advantages

for dykes, flood protection and barrier dams, floating

embankments, counterweight fills and retaining structures

with and without geosynthetic reinforcement. Moreover, it

has proved successful for roads, highways, railways and 

all kinds of earthworks and reinforced retaining structures.

Compared with conventional spot checking CCC provides

•   increased safety

•   improved serviceability

•   improved seismic response

•   increased lifetime

•   reduced costs and maintenance

•   reduced construction time

•   detailed monitoring of the top zone (2.5 m) of 

existing dykes or dams (see Figure 10).

Furthermore, CCC increases the installation survivabil-

ity of geosynthetics placed in multi-layered structures. The

uniformly compacted, smooth surface of each fill shift

reduces the risk of wrinkles or punch-through in the

Display unit

Drum

Sensor 

Off-centrerotatingmass

Processor 

 Auxiliary sensor 

Figure 72. Principle of roller-integrated continuous

compaction control (CCC) and compaction optimisation

0.000 0.005 0.010 0.015 0.020 0.025 0.030 0.0351000

800

600

400

200

0

Vibroflotation points

Omega

Limit for sufficientcompaction

Excavation level

Vibroflotation area  Area without

vibroflotation

Figure 73. Results of roller-integrated continuous compaction

control (CCC) on a 1.5 m thick fill on top of subsoil

improved by deep vibroflotation. The positions of the stone

columns are still clearly visible on the CCC diagram.

Omega dimensionless value of compaction degree or soil

stiffness respectively

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geosynthetic; also, local overcompaction of the fill layers

(and hence stress constraints in the geosynthetic inclusions

or cover) is avoided.

Intensive and uniform compaction of geosynthetic-

reinforced dykes, dams, barrier fills and other retaining

structures improves their earthquake resistance signifi-

cantly. If the layers are not properly compacted they will

 be more damaged during stronger seismic events. A rigid facing of geosynthetic-reinforced retaining walls makes

higher compaction of the outer fill zone possible. Conse-

quently, such structures suffer less damage. Over-intensive

compaction, should be avoided, though, because it creates

excessive lateral earth pressures behind rigid facings, or 

moves flexible elements outwards. CCC is an excellent

tool for optimising compaction intensity, depending on the

structural requirements.

The increase in lifetime and serviceability of structures

is especially important for road and highway pavements

and for railways, but also for dykes and dams, and for 

 barrier fills and other retaining structures. Weak spots in

dykes or dams may promote seepage, inner erosion or 

 piping: hence not only the absolute degree of compaction

 but also its uniformity is essential. This can be best

checked by CCC, which is a significant advance over the

statistical quality control used hitherto. Until now the

various selection procedures for spot checking have been

grid pattern, random selection, subjective selection, and 

subjective selection using existing criteria.

The CCC method also involves new statistical criteria,

 because it records the uniformity of the compacted layer 

as well as the increase in degree of compaction during

subsequent roller passes. The relevant parameters are the

mean value, maximum and minimum values, standard deviation, and increase of the compaction values. More-

over, the minimum quantile can be used as soon as

sufficient experience with CCC is gained on the particular 

construction site.

The statistical parameters used hitherto do not allow

assessment of the distribution of control data within a

section: The plots in Figure   74   have the same mean

maximum and minimum values, and standard deviation,

 but they exhibit different qualities. Figure 74a shows only

one limited weak zone, which can easily be improved,

whereas the area in Figure   74 b requires comprehensive

measures to achieve sufficient and homogeneous quality.

To judge the distribution of control values statistically

within a defined area, the statistical methods of vario-

graphy and kriging can be used. These methodologies

were originally developed in order to interpret geophysical

data, and seem to be a useful tool to improve the quality

assurance of compacted layers, and hence also of geosyn-

thetic–soil structures.

A reliable interpretation of the CCC data is possible

only if the operating conditions of the vibratory roller 

drum are taken into consideration (see RVS 8.S.02.6 (FSV

2005)). The significant operating conditions depend on

the roller and soil data, and on the interaction between

roller and soil (or other granular material) (Adam 1996;Brandl and Adam 1997).

Roller compaction technologies have been improved 

significantly over the last 15 years, and they now provide

a wide range of possibilities for selecting the most suitable

roller for a particular purpose. A further development is

the automatically controlled VARIOCONTROL roller,

whereby the direction of excitation is controlled automati-

cally by using defined control criteria. VARIOCONTROL

compaction provides uniform compaction, fewer roller 

 passes, improved compaction in both deeper layers and on

the surface, and reduction of lateral vibrations, for exam-

 ple when operating close to sensitive structures.

Smaller compaction equipment (plate vibrators, etc.)

also exists that facilitates CCC for backfills under con-

fined site conditions, or for slender geosynthetic–soil

structures.

To sum up, roller-integrated CCC represents a signifi-

cant improvement for high-quality management systems.

Compaction control is integrated in the compaction pro-

cess, and data are provided over all the compacted area.

Because of the outstanding advantages of CCC, this

technology should be used for the compaction of soil

structures as much as possible, especially in the case of 

geosynthetic reinforcement.

To a certain extent, compaction control can be alsoachieved by the spectral analysis surface wave method 

(SASW) or continuous surface wave technique (CSW).

(a)

(b)

Figure 74. Control plots of compacted areas with the same

conventional statistical characteristics of compaction

(schematic). The strips indicate the roller lanes, controlled

with CCC. The different shadings show different compaction

degrees. Statistical evaluation can be improved with

‘variography’ or ‘kriging’

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for fixing the absorber pipes and for overlapping the

geotextiles

•   absorber pipes of linear polyethylene with copolymer 

octane, outer diameter 25 mm, wall thickness

3.5 mm, exhibiting high flexibility and stability,

suitability for bending with small radius but small

rebound forces, and stable long-term behaviour.

The geotextiles were fixed to the primary tunnel lining

(shotcrete) with common nails and, additionally, the

absorber pipes were fixed with pipe clamps to avoid 

deformation during concreting of the secondary lining

(Figure   79). Four strips of energy geocomposite were

connected to one thermo-active unit with one manifold for 

absorber entry and another one for absorber return. There-

fore the main attention during filling the absorber system

with antifreeze must be paid to complete filling and 

degassing of all absorber pipes. Preliminary tests had been

conducted to find a suitable technique for filling the

absorber system, as there is no possibility for degassing at

the top(s) of the absorber system. The passage of thecollecting absorber pipes requires special measures, as

illustrated in the examples of Figures   80   and   81.   In

 principle, there are four options: plastic screws, flexible

hose, niches, and reinforcement crossing.

The pipes have to withstand the outer pressure of 

casting the concrete: therefore they must be kept under 

inner pressure (compressed air,   pi > 2.0 bar) during con-

creting. The formwork transport equipment for the sec-

ondary tunnel lining needs special openings for the pipe

 passages. The absorber pipes coming from the individual

geosynthetic strips have to be connected to the collecting

 pipes by special T-shaped elements (Figure   82). The

collector pipes should run as far as possible between the

inner and outer reinforcement of the tunnel lining (Figure

83). After finishing all welding work, the entire pipe

system has to undergo detailed tightness testing (Figure

84). The tightness of the pipe system must be checked 

after each construction phase: that is, before placing the

reinforcement, after welding the bars, etc.

6.5. Test results

The energy extracted from the ground was transferred to a

heat pump and then to a radiator that emitted the produced 

heat to the air in the tunnel. Thus the efficiency of the

thermo-active system could be measured. For parametricstudies, different operating features were investigated. The

measurements comprised (Markiewicz 2004)

16.00 m

Shotcrete

Inner lining

Slidingmembrane

     1 .     8

     4     m

        1 .        8

        4       m

9.94 m

7.00

 Absorber pipes  Absorber pipes

Energy geotextile Energy geotextile

Watertight connection

Manifold to heat pump

2%

Figure 78. Cross-section of the test site

Figure 79. Installation of energy geotextiles at the test plant

LT22-Bierhauselberg. Pipe loops clearly visible

Table 4. Technical data of the energy geotextile

Property Value

Mass per unit area (g/m2) 285

Tensile strength (kN/m) 21.5

Elongation at maximum load (%) 100 (MD)

40 (CD)

Static puncture resistance (CBR test) (N) 3300

Cone drop test (hole diameter) (mm) 17

Permeability (vertical) (l/m2 s) 70 l/m2s

Opening size,   O90   ( m) 95

Thickness (2 kPa) (mm) 2.5

382   Brandl 

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•   the temperature of the brine solution fluid in the

absorbers (entry, exit)

•   the temperature of ground/groundwater and tunnel

(air, lining)

•   the heat production of the radiator 

•   the flow velocities in the heating circuit and absorber 

fluid circuit

•   the electric current consumption of the heat pump

and circulating pumps•   photographic documentation of the surface of the

secondary tunnel lining, with thermal imaging

Inner lining

Inner lining

To energygeotextile

To energygeotextile

To heatpump

To heat

pump

(a)

(b)

Figure 80. Watertight passage of the collecting absorber

pipes through inner lining with (a) screws or (b) flexible hose

Figure 81. Pipe passage through the secondary (inner) tunnel

lining. Detail with criss-cross reinforcement

Figure 82. T-shaped elements to connect the absorber pipe

from the individual strips of energy geosynthetics to the

collector pipes

Figure 83. Collector pipe running between the inner and

outer reinforcement of the tunnel lining

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With regard to thermo-active geocomposites, the fol-

lowing innovative structure was designed: primary lining

 – energy geocomposite – secondary lining. This geocom-

 posite consists of three elements: a nonwoven geotextile,

absorber pipes and a sliding membrane. Geotextiles with a

high in-plane drainage capacity (high transmissivity) can

fulfil the drainage function of a structured geomembrane,

if only moderate water ingress is expected. Figure   86shows an installation detail for such ‘energy geocompo-

sites’.

Sliding membranes of 0.2 mm thickness, commonly

used on construction sites, have proved suitable. Small

leaks would not affect their serviceability, because they

serve not for waterproofing but for separation, with a

smooth interface between geocomposite and waterproof 

concrete. This reduces stress constraints within the tunnel

lining, and minimises the transfer of shear forces on the

waterproof concrete (according to the ‘white tank’ tech-

nology).

6.7. Summary

Thermo-active ground structures (energy foundations,

retaining walls, tunnels, etc.) and also energy wells are

a promising innovation in sustainable and clean energy

consumption. A significant advantage of such systems

is that they are installed within elements that are

already needed for statical/structural or geotechnical

reasons: hence no separate or additional structural or 

hydraulic measures are required. Foundations, walls

(below and above ground) or tunnel linings can be used 

directly for the installation of absorber pipes for heat

exchange.

Energy tunnels using geosynthetics fitted with absorber  pipes are a key improvement over conventional geother-

mal methods such as deep borehole heat exchangers or 

near-surface earth collector systems. Comparative large-

scale tests and site experience have shown that compo-

sites, consisting of nonwoven geotextiles fitted with

absorber pipes provide the best results. Usually, a tempera-

ture difference of only  ˜T  ¼ 28C between absorber fluid 

inflow and return flow from the primary circuit is

sufficient for an economical operation of the energy

system. Consequently, such geothermal systems represent

low-temperature systems. Experience has shown that the

electricity required for operating the entire system com-monly varies between 20% and 30% of the total energy

output. If no heat pump is necessary (e.g. for free cooling)

this value drops to 1–3% for merely operating a circula-

tion pump.

7. FURTHER APPLICATIONS OFGEOSYNTHETICS

In addition to the previous chapters some special applica-

tions of geosynthetics are mentioned below

7.1. Energy piles

Energy piles represent an innovative, environmentallyfriendly technology for heating or cooling buildings

(Brandl 1998, 2006b). In Austria, nearly 70 000 energy

 piles have been installed since the 1980s. They provide

substantial long-term cost savings and minimised mainte-

nance. If piles already required for geotechnical or structur-

al reasons, are equipped with absorber pipes, they work 

simultaneously as bearing elements and as heat exchangers.

Moreover, they may be combined with diaphragm walls,

rafts or basement walls (e.g. Figure 87). However, in very

soft soil the diameter of in situ cast concrete piles may be

constricted. This can be avoided by wrapping the reinforce-

ment cage of the piles with highly stretchable nonwoven

geotextiles (welded). Bulging can be minimised as well.

When casting the fresh concrete, the lateral pressure of the

fluid stretches the geotextile, so that a sufficient concrete

cover of the reinforcement is attained. This technique

20 20 20

40

 Absorber pipes25 mm dia.

Slidinggeomembrane

Nonwovengeotextile

Figure 86. Installation detail for overlapping thermo-active

‘energy composites’: Nonwoven geotextile (with high

transmissivity), absorber pipes, sliding geomembrane

Thermo-activeconcrete slabs

Secondarythermal circuit

Thermo-activeconcrete slabs

Energy piles

Primarythermal circuit

Energydiaphragm

walls

Figure 87. Cross-section through a building with geothermal

cooling and heating; energy piles and energy diaphragm

walls in very soft ground

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 provides the required pile integrity, and hence the designed 

heat transfer from and into the ground.

7.2. Rockfall shelters

Roads and railways along steep rock slopes need shelters

if protective fences are not sufficient to withstand severe

impacts. Large-scale blocks of rock falling from great

height may destroy even prestressed reinforced concretestructures if they hit them directly (Figure   88). Therefore

shelters under high dynamic load should be covered by

sloped soil fill, preferably reinforced with geosynthetics

(Figure   89): thus the blocks can roll over the structure

without damage.

7.3. Sandwich reinforcement of embankments and

barrier dams

The stability and deformation behaviour of high embank-

ments and barrier dams along steep or unstable slopes, or 

in seismic areas, can be significantly improved by geosyn-

thetic sandwich reinforcement. The scheme in Figure   90

shows that the reinforced zones should preferably be

 placed at levels where benches are designed. The top

reinforcement is recommended if the embankment carries

traffic routes. It serves for better load distribution, and 

reduces differential settlements. Moreover, sandwich rein-

forcement provides a high earthquake resistance.

7.4. Glacier melting

Glacier melting progressively creates unstable slopes,

when ground that had been frozen into great depth is

thawing. High pore-water pressures and seepage reduce

local and global slope stability, and trigger debris flows,

multiple slips or deep-reaching block sliding. Moreover,

glacier melting becomes a problem in tourist regions,

especially where ‘Alpine’ skiing dominates. Therefore

several precautionary measures and contingency plans

have been developed in Austria. A special application of 

geosynthetics is the local cover of exposed glacier areas

and zones around lift columns with white, nonwoven

geotextiles (Figure   91). Lift columns are frequently

founded or anchored in creeping ice that undergoes in-

creasing differential movements. A local reduction of 

glacier melting by geotextile cover prolongs the intervals

of column readjustment.

Several in situ tests were performed to find a method to

reduce glacier melting (Figure 92). Measurements showed 

that special geotextiles provided the optimal results, with

a reduction of about 60%. The main characteristics of themost suitable protective geosynthetics (found from com-

 parative in-situ tests) are listed in Table 5.

7.5. Beaver barriers

An increasing population of beavers also represents an

increasing danger to dykes and flood protection dams.

Therefore field tests with several geosynthetics as beaver 

 barriers were conducted in cooperation with zoologists

(Figure  93). The tests showed that the beavers eventually

 broke through all the geosynthetics, but not through the

metal wire mesh with synthetic cover. Such wire meshes

have been installed in Austria since 2008 along criticalsections through and beneath dykes and dams. At the toe

of the dam they are fixed in a trench filled with concrete

Figure 88. Concrete shelter damaged by heavy rockfall

  O  r  i  g   i

  n  a  l   t  e

  r  r  a  i  n

  O  r  i  g   i

  n  a  l   t  e

  r  r  a  i  n

 Anchor wall(on filter concrete)

Reinforced withgeosynthetics

Finalterrain

  2  :  3

Sockets Dolomite(mylonitic)

 Al 

 A

1000 kN25–35 m

Prestressedanchors

Variablelength

1 : 5

  2  :  3

140 m4.7m

Figure 89. Reinforced concrete shelter (gallery) against

rockfall, debris flow and avalanches, covered by protective fill

reinforced with geosynthetics;   Ar   is anchor force

Interlocking(toothed)

Top reinforcement for roads, highways or heavy buildings

(Possibly interlayersin between)

Reinforced interlayers( 3–6 m)h

Not reinforced

Natural ground(unstable)

Figure 90. Sandwich reinforcement of high embankments

and barrier dams along steep or unstable slopes and in

seismic zones. Scheme with steepened fill slope; not to scale

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(Figure 94); at the crest they are covered by the layers of 

the road structure.

7.6. Geosynthetics for in-situ groundwater cleaning

Permeable reactive walls or funnel and gate systems have

 proved suitable for the in situ cleaning of contaminated 

groundwater (Figure   95). The contaminant plume then

flows through a straight or curved wall, or is directed to agate. Groundwater cleaning in the reactive wall or gate is

 performed site-specifically, whereby physical, chemical,

and microbiological measures are possible. Several sys-

tems contain exchangeable geosynthetic filter panels, and 

also geotextiles to envelope special (granular) reactive

material.

Figure 91. Local geotextile cover of glaciers to reduce

melting of particular zones

Without measures

TenCate Topex GLS

Water injection

Snow compaction

300

250

200

150

100

50Moltenthicknessofice/snow

(cm)

60% reductionof snow melting

July 2005 September 2005

Time

Figure 92. Reduction of glacier melting by placing a special

nonwoven geotextile cover; results of in situ tests

(a)

(b)

(c)

Figure 93. (a, b) Geosynthetics and (c) wire mesh to prevent

damage to flood protection dams caused by beavers. Field

tests in cooperation with zoologists

Table 5. Geotextile data to reduce glacier melting

Property Value

Geotextile Nonwoven

 polypropylene

Mass per unit area (g/m2) 340

Thickness (at 2 kPa) (mm) 2.0

UV resistance (residual strength) (%)   . 75

Tensile strength (EN ISO 10319) (kN/m) 23

Opening size (EN ISO 10319) (mm) 0.06

Thermal resistance (m2K/W) 0.0593

Water permeability in plane (m3/m s) 2 3 10 –6

Water penetration resistance (cm)   . 8

Light reflection capacity (wave length 500 mm) (%) 78

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7.7. Contaminant-absorbent geosynthetics

Contaminant-absorbent geosynthetics are not only used in

reactive walls to clean groundwater but also for surface

measures. In both cases nonwoven geotextiles are pre-

ferred for environmental protection. Whereas reactive

walls are installed more for long-term purposes, surface

measures serve for emergency action (Figures  96  and  97).

Absorbent geosynthetics used for surface protection or 

cleaning may be socks, pillows, mats, rolls or oil booms.

Furthermore, oil skimmers may absorb oil and fuel in

canals, bilges, sumps, manholes, drains and tanks.

Absorbent nonwoven geotextiles consist of polypropy-

lene. Their costs are clearly below those of granular 

absorbers, and their absorbent capacity is significantly

higher. The used (and hence contaminated) absorbent

geosynthetics are most suitable for waste incineration and 

thermal utilisation. The main characteristics of the hitherto

optimised product are

•   production technology: meltblown

•   fibre fineness: 5– 8  m

•   bonding type: thermally calender or ultrasonic

 bonding

•   engraving: point bonded, flat bonded 

•   mass per unit area: 15– 400 g/m2.

Oil booms and skimmers consist of polypropylene

geotextile strips covered by a nylon net of high tensile

strength.

Figure 94. Installation of nets against beaver activity along

flood protection dams

Ground plans

Contaminationsource

Contaminationsource

Reactive wall

Cleanedgroundwater 

Cleanedgroundwater 

Cleanedgroundwater 

Cleanedgroundwater 

Contamination plume Contamination plume

GW flow direction GW flow direction

Funnel

Unsaturated zoneUnsaturated zone

 Aquifer   Aquifer 

 Aquitard  Aquitard

k k f r -Wall -Aquifer  

(a)

k k k k 

r f 

r f 

-Gate -Aquifer  -Funnel -Aquifer  

(b)

Gate

Cross-sections

Figure 95. In situ groundwater cleaning with (a) permeable reactive walls or (b) ‘funnel and gate’ system

Figure 96. Placing of geosynthetic mats to absorb

contaminating liquids on a river

388   Brandl 

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8. FINAL REMARKS AND

CONCLUSIONSGeosynthetics engineering offers a broad range of 

applications for damage prevention, natural disaster 

mitigation, rehabilitation measures, and environmental

 protection. Coastal protection, erosion protection, seismic

aspects, sinkhole bridging and hazardous waste contain-

ment, for instance, are topics that have been treated in

numerous publications. Therefore they are not discussed 

in this Giroud Lecture, although referred to in most

sections.

The selected geosynthetic applications emphasize the

close links between theory and practice, and between

geosynthetics and geotechnical engineering. Design byfunction, or semi-empirical design with calculated risk 

(interactive design), combined with the observational

method, represents an essential feature of geosynthetics

application. For geosynthetic–soil systems in critical

areas (unstable slopes, seismic zones, etc.) contingency

 plans should exist (e. g. possibilities of in-time strength-

ening).

Because of their flexibility, geosynthetic–soil struc-

tures have clear advantages over rigid structures with

‘classical’ construction materials: higher impact absorp-

tion of rockfalls or avalanches, lower sensitivity to

earthquake and creeping slopes, and rapid installation

(for flood defence, etc.). Ecological advantages can also

 be found when comparing the total energy consumption

for the component and final products, for transportation

from the manufacturer to the construction site, and for 

installation.

ACKNOWLEDGEMENTS

This lecture was presented at the 9th International Con-

ference on Geosynthetics, in 2010, in Guaruja, Brazil. The

author wishes to thank the publishers (Brazilian Chapter 

of the International Geosynthetics Society) and the organi-

zers (E. M. Palmeira, D. M. Vidal, A. S. J. F. Sayao and M. Ehrlich) for permission to publish this paper in an

updated and extended version.

NOTATIONS

Basic SI units are given in parentheses.

 A   area (m2)

 A9   corrected area (m2)

c   cohesion (N/m2)

cr    residual cohesion (at larger shear deformations)(N/m2)

 DPr    degree of compaction (percentage of Proctor 

density)

 E cr    creep pressure (N/m2)

 E kin   kinetic energy (J)

 E  pot   potential energy (J)

 E v   deformation modulus (N/m2)

 F    safety factor (dimensionless)

 F    horizontal shear force (N)

FS factor of safety (dimensionless)

G    shear modulus (N/m2)

 H    height (m)

h   height (m)

icrit   critical hydraulic gradient (dimensionless)

 ss   self-settlement (mm)

    performance factor (dimensionless)

    slope angle (degrees)

r   density (g/m3)

  friction angle (degrees)

r    residual shear angle (degrees)

design   design shear angle (degrees)

 peak    peak value in the curve    versus shear 

deformation (degrees)

REFERENCES

Adam, D. (1996).   Continuous Compaction Control with Vibratory

 Rollers. Doctoral thesis, Vienna University of Technology, Vienna,

Austria (in German).

Akkerman, G. J., Bernardini, P., van der Meer, J., Verheij, H. & van

Hoven, A. (2007). Field tests on sea defences subject to wave

overtopping.   Coastal Structures 2007: Proceedings of the 5th

Coastal Structures International Conference, CSt07-2, Venice, Italy,

Franco, L., Tomasicchio, G. R. & Lamberti, A., Editors, World 

Scientific, Singapore, vol. 1, pp. 657–668.

Blovsky, S. (2002).   Bewehrungsmoglichkeiten mit Geokunststoffen.

Doctoral thesis, Institute for Soil Mechanics and Geotechnical

Engineering, Vienna University of Technology, Vienna, Austria (in

German).

Brandl, H. (1979). Design parameters for weak rocks. Panel discussion.

 Proceedings of the 7th European Conference on Soil Mechanics

and Foundation Engineering , Brighton, UK, A.A. Balkema,

Rotterdam, the Netherlands, vol. 4, pp. 63–67.

Brandl, H. (1980). Theory of earth pressure and creep pressure of sliding

slopes with practical examples.   Proceedings of the 6th Danube-

 European Conference on Soil Mechanics and Foundation Engineer-

ing , Varna, Bulgaria.

Brandl, H. (1998). Energy piles and diaphragm walls for heat transfer 

from and into the ground.   Proceedings of the 3rd International 

Seminar on Deep Foundations on Bored and Auger Piles, Ghent,

Belgium, Van Impe, W. F., Editor, A.A. Balkema, Rotterdam, the

 Netherlands, pp. 37–60.

Brandl, H. (2006a). Bridges or embankments in creeping slopes?

 Darmstadt Geotechnics, Technical University Darmstadt, Germany,vol. 11, pp. 73–92.

Brandl, H. (2006b). Energy foundations and other thermo-active ground 

structures. Rankine Lecture.   Geotechnique,  56, No. 1, 79–122.

Figure 97. Oil booms of absorbent geosynthetics to contain

and absorb oil spills on water

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