ORIGINAL PAPER
Re-assessing a soil nailing design in heavily weathered graniteafter a strong earthquake
Sergio A. Villalobos • Paulo L. Orostegui •
Felipe A. Villalobos
Received: 26 August 2011 / Accepted: 18 January 2013
� Springer-Verlag Berlin Heidelberg 2013
Abstract Soil nailing is increasingly adopted in poten-
tially unstable slopes and steep excavations in Concepcion,
Chile. The paper presents a study of soil nailing design and
construction in Maicillo, a heavily weathered granite. The
stability of the soil nailed wall was analysed using a limit
equilibrium block method. The lowest factor of safety was
obtained when the nailed wall was 8 m high but without
the bottom row of nails. This was followed by the global
seismic case using an acceleration factor kh = 0.15. The
occurrence of a strong earthquake (magnitude 8.8) 2 weeks
after completion of the projected allowed a re-assessment
of the soil nailing design. Despite registering a maximum
acceleration of 0.63 g near the project site, the nailed wall
did not present any damage, probably due to the use of
undrained shear strength parameters.
Keywords Nailed wall � Heavily weathered granite
(Maicillo) � Limit equilibrium method � Stability analysis �Strong subduction earthquake
Resume Le clouage des sols est une technique de plus en
plus adoptee pour les pentes potentiellement instables et les
excavations a pentes raides a Concepcion, au Chili.
L’article presente une etude de dimensionnement et mise
en œuvre d’un clouage de sol dans une zone de granite tres
altere, nomme Maicillo. La stabilite de la paroi de sol cloue
a ete analysee en utilisant une methode d’equilibre limite.
Le plus faible coefficient de securite a ete obtenu lorsque la
paroi clouee faisait 8 m de hauteur, mais sans la rangee
inferieure de clous. Cette situation etait suivie par le cas
sous sollicitation sismique avec un facteur sismique
kh = 0,15. Un seisme de forte magnitude (M = 8,8) deux
semaines apres la realisation de l’ouvrage a permis de
reevaluer le dimensionnement de la paroi clouee. En depit
de l’enregistrement d’une acceleration maximale de 0,63 g
a proximite du site de projet, la paroi clouee ne presentait
aucun dommage, probablement en raison de l’utilisation de
parametres de resistance au cisaillement non draine.
Mots cles Paroi clouee � Granite fortement altere
(Maicillo) � Methode des equilibres limites � Analyse de
stabilite � Fort seisme de subduction
Introduction
The reduction of available urban space and the hilly
topography in the Greater Concepcion area, Chile, increase
the need for building in challenging conditions. Housing
developments and motorway projects often involve the
re-profiling of slopes, or simply that they are cut vertically
to create flat spaces. This can result in unstable slopes
which need to be retained usually with traditional gravity
or cantilever T or L-shaped reinforced concrete walls.
However, rigid wall solutions are expensive, time con-
suming and generally restricted to heights of 5 m due to the
risk of overturning.
S. A. Villalobos
Geology and Geomechanics, New Mine Level Project,
CODELCO, Coya Camp 267, Machalı, Rancagua, Chile
e-mail: [email protected]
P. L. Orostegui
OITEC Engineering, Surveying and Geotechnics,
Lincoyan 444, Of. 309, Concepcion, Chile
e-mail: [email protected]
F. A. Villalobos (&)
Department of Civil Engineering, Catholic University
of Concepcion, Alonso de Ribera 2850, Concepcion, Chile
e-mail: [email protected]
123
Bull Eng Geol Environ
DOI 10.1007/s10064-013-0466-7
Alternatively, soil reinforcement construction tech-
niques may be appropriate. These have been increasingly
applied worldwide for at least 30 years and can be classi-
fied broadly into two categories:
1. Earth reinforcement structures require the use of
reinforcing elements and a fill material. The former
can be a type of geotextile, geogrid, metal rods or
meshes and the latter can be well compacted sand.
2. The second category does not incorporate a fill
material, but introduces reinforcement elements, such
as piles, micropiles, anchors and nails, within the
natural soil. For this reason, it is generally called
in situ soil reinforcement.
This paper concerns the latter, i.e. nails in residual
soils. Soil nailing is a solution to stabilise slopes or ver-
tical excavations such as those depicted in Fig. 1. The soil
nailing construction technique consists of driving or
drilling and subsequently grouting under high pressure
around inclined bars as the excavation or re-profiling
progresses downwards. Bars are normally inclined
downwards at between 10 and 30� and spaced horizon-
tally and vertically at between 1 and 2 m, depending on
the specific calculations. The exposed surface is rein-
forced with steel meshes and sprayed with shotcrete to
create a flexible structural facing with a thickness of
between 50 and 250 mm. A drain system is provided for
the evacuation of water behind the shotcreted wall. The
reinforcing bars transmit tension loads, although they can
resist limited shear and bending loads which are usually
neglected (Guilloux et al. 1983). The tensile load distri-
bution has been measured by Guilloux et al. (1983) who
found that it varies not only along the bars but also with
the depth of the bar. The maximum tensile load in the bar
is just behind the facing at the toe of the slope or exca-
vation, and gradually separates from the facing upwards.
This creates a parabolic distribution of maximum tensile
stresses, which tends to follow the potential failure
surface.
Soil nailing is increasingly being chosen around
Concepcion, in projects where residual soils are found.
However, few available publications are related to soil
nailing applications in these soils (heavily weathered
granite) and in a highly active tectonic area. This work
does not address the deformation and displacement of nails
or walls. Instead, the stability problem is solved using the
method of limit equilibrium of forces, which is normally
used in the Chilean practice.
Global stability analysis
A global stability analysis considers the development of
failure surfaces in the ground, which may or may not
intersect the nails. Ideally, a design should consider the
intersection of nails with the potential failure surface in
order to increase the soil resistance and contribute to the
ground stability. Stocker et al. (1979) carried out several
trials in nailed walls in Germany to study failure mecha-
nisms. They found that the failure surface is not actually
curved, but resembles more two linear parts, as shown in
Fig. 2. Subsequently, Gaessler and Gudehus (1981) con-
firmed the bi-linear shape of the failure mechanism in
laboratory tests of scaled soil nailing models. Figure 2
shows an application of the model proposed by Stocker
et al. (1979), where a bi-linear failure surface generates two
sliding blocks. This analysis method is also known as the
sliding block method.
The equilibrium calculations consider two rigid sliding
blocks separated by a vertical line bd; one is reinforced
with nails and the other is an active wedge. The angle h2
between bc and the horizontal is assumed to be equal to
h2 = p/4 ? //2. Then, this failure surface along bc cor-
responds to an active Rankine state. The inclination angle
h1 between ab and the horizontal is determined by iteration
until a minimum value of the safety factor is reached. The
total static active thrust Ea(2–1) due to the lateral earth
pressure distribution acting on the fictitious wall bd can be
expressed as:
Eað2�1Þ ¼1
2cL2
bd þ qLbd
� �ka � 2cLbd
ffiffiffiffiffika
pð1Þ
Fig. 1 Typical soil nailing
applications, reinforcing
a natural slopes and
b excavations
S. A. Villalobos et al.
123
where c is the unit weight of the soil, q is a uniformly
distributed overburden, c is the cohesion and ka is the
Coulomb active lateral earth pressure coefficient. To resist
Ea(2–1) the soil will mobilise its shear strength along ab and
bc. It is assumed that no shear is developed along bd, i.e.
both blocks move together. The Coulomb failure criterion
is used to evaluate the maximum soil shear strength
mobilised along the failure surface. A global factor of
safety is defined as the ratio between the sum of resisting
forces and the sum of applied forces. The resisting forces
C1 and C2 are obtained from the soil shear strength along
the bi-linear failure surface.
C ¼ C1 þ C2 ¼ cLab þ N1 tan /þ cLbc þ N2 tan / ð2Þ
where / is the angle of friction in plane conditions;
however, it is common practice to use / values determined
from drained or undrained triaxial tests, instead of, for
example, direct shear tests. The normal force Ni is given
by:
Ni ¼ ðWi þ QiÞcos hi
cos /ð3Þ
where the subindex i = 1 ? reinforced block and i = 2 ?active wedge, Wi is the weight of the block and wedge, Qi
is the overburden and hi is the inclination angle. The
tension force contribution of the n passive nails can be
expressed as:
T ¼Xn
i¼1
Tpi ð4Þ
where Tpi is the resistance force added by the nails crossing
the failure surface. The tension capacity is a structural
property of the nail material and the pull-out capacity
results from the interaction between the nail, the grouting
and the surrounding soil.
The global factor of safety FSG is given by the ratio
between the resistant forces Ci, Ni and T and the applied
forces Wi, Qi, Ea(2–1) and Hi.
FSG ¼Ci þ Ni þ T
Wi þ Qi þ Eað2�1Þ þ Hið5Þ
The horizontal force Hi applied in the gravity centre of
both blocks represents a dynamic force in a pseudo-static
form. In earthquake geotechnical engineering practice, H is
represented as product between the weight (in this case the
weight of the rigid block) and an acceleration factor (kh)
representing the maximum horizontal acceleration
component of the earthquake.
H ¼ khW ¼ amaxh
gW ð6Þ
where amaxh is the maximum horizontal acceleration, usually
determined as a reasonable value for the geographical zone
under study, and g is the acceleration of gravity. As in the
analysis of other retaining structures, the resistance against
sliding along Lab and the bearing capacity of the foundation
soil under the soil nailing are also evaluated.
Case study
Geological and geotechnical background
The geological unit present in the project site of Quinta
Junge corresponds to a heavily weathered igneous intrusive
rock which forms part of the coastal batholith mountain
(see Fig. 3). This Palaeozoic granitic rock is between 250
and 570 million years old and is the result of the tectonic
activity caused by the subduction of the Nazca plate under
the South American plate. Weathering has destroyed the
bonding between the mineral grains and has transformed a
strong rock into crumbly lumps—the residual soil being
known locally as Maicillo. This type of material is very
complex to analyse as it is difficult to determine whether it
will behave as a rock or soil, or a combination of both.
When dry, the material has a high resistance attributable to
the remaining bonding of the granite rock. When it is wet/
Fig. 2 Global stability analysis
using sliding block method
showing, a forces acting on the
reinforced soil block and
b polygons of forces
Soil nailing in granite
123
saturated, the silt particles dominate and an undrained
behaviour prevails, which can lead to the sliding of blocks
along fracture planes in the original rock.
Figure 4a shows the blocks of the weathered granite
obtained for sampling and testing. It can be seen that the
material is composed of coarse grains, which may disin-
tegrate when handled (Fig. 4b).
Table 1 summarises the values of the main geotechnical
parameters of the Maicillo which consists of 30 % fine and
medium gravel, 55 % sand and 15 % silt of low plasticity.
The yellowish brown material classifies as a silty sand SM,
according to the USCS. It is worth mentioning that this
classification is obtained from testing a granular material
(Fig. 4b). However, intact samples were tested in a direct
shearbox apparatus. Only slight remoulding around the
corners was required, as can be seen in Fig. 4c.
The construction problem
The original seven storey block of flats in Quinta Junge
were constructed in an excavation supported by stone
masonry walls and shotcrete. The enabling works for a
second block of flats adjacent to the first required an 8 m
deep excavation with a back face at 70� cut into a 35� hill
slope, again supported by stone masonry walls and shot-
crete. However, during the excavation of the first 4 m
below a masonry wall, a triangular sliding failure occurred
(Fig. 5). The reduction of resistance as a consequence of
the excavation induced sliding along pre-existing planes of
failure in the weathered Maicillo. A common solution in
the zone, mainly for sandy soils, is anchored soldier pile
walls. However, this solution was not possible as the steel
soldier piles cannot be driven into Maicillo. Finally, it was
decided to adopt a temporary soil nailing system until the
building was completed and the intervening space
backfilled.
Soil nailing design
The soil nailing design was undertaken using the computer
program GGU-Stability (2008) which calculated the con-
ditions of limit equilibrium for the failure mechanism
presented in Fig. 2. The inputs used in the analysis con-
sidered a 0.15 m thick shotcreted wall with four rows of
nails. Table 2 summarises the parameter values used in the
analysis and design of the soil nailing. The 2 m high stone
masonry wall was treated as a material with a unit weight
of 25 kN/m3 as shown in Fig. 6.
The nails used were high resistance thread steel bars of
25 mm nominal diameter with a minimum failure resis-
tance of 630 MPa and a yield stress fy between 420 and
580 MPa. This results in an allowable tension capacity of
180 kN, defined as the 90 % of the yield stress for a bar
section Ab, Ta = 0.9fyAb. The soil–nail interface shear
strength rs was estimated from a chart for sands proposed
by Bustamante and Doix (1985), where rs is correlated with
an SPT value of 30 blows/foot. The analysis assumed an
effective angle of friction /0 = 30� and a cohesion
c = 10 kPa based on the work of Ruız (2002) who carried
out undrained triaxial tests on Maicillo samples. These
values were mobilised under an axial deformation of 20 %,
representing a sliding condition. Yin et al. (2009) presented
cohesion and friction angle values from consolidated
drained triaxial tests carried out in a loose and completely
decomposed granite from Hong Kong. For Sr = 50 %,
c = 44 kPa and / = 31.9� and for Sr = 98 %, c = 5 kPa
and /0 = 34.6� for an axial deformation of 20 %.
Undrained shear strength parameter values were used to
allow for saturated and rapid loading conditions as a worse
case scenario.
Design of the construction sequence
The soil nailing construction began with the re-profiling of
the slope face to 85� for the installation of an electro
welded mesh ACMA C257 with a grid of 150 9 150 mm
of 7 mm nominal diameter steel wires with a yieldingFig. 3 General geological map of Concepcion (Poblete and Dobry
1968)
S. A. Villalobos et al.
123
stress of 500 MPa. Subsequently, nails were installed fol-
lowed by the addition of four 1.2 m long and 12 mm
diameter steel bars centred in a cross form around the nail
head. The purpose of the steel bars was to transmit the
loads taken by the nails to the shotcrete wall and reduce the
risk of flexural/shear failure. The steel bars were sand-
wiched by a second electro welded mesh ACMA C257,
resulting in a total of 514 mm2/m2 of steel mesh. Shotcrete
was injected to form a 30 mm thick layer between the soil
and the inner mesh and between the outer mesh and the
exterior wall face, resulting in a total wall thickness of
0.15 m. Finally, a 0.2 m square bearing plate and a nut
were installed.
The construction sequence is shown in Fig. 6 for the
static cases. Although the groundwater table may reach the
inner end of the nails, it is significantly further back than
the potential sliding blocks. The anticipated sliding blocks
would not extend back as far as the swimming pool, some
Fig. 4 a Heavily weathered
granite rock blocks in Quinta
Junge, b Maicillo after sieving
showing coarse and fine
components and c sample
preparation in the shear box
Soil nailing in granite
123
7.5 m from the nailed wall, except when seismic forces are
involved, as described below.
The installation of the nails was carried out using a
Comacchio MC 600 drill rig with a tricone bit. Air injected
with eight bars (800 kPa) pressure was used as a flush.
During the drilling operations, an electro-welded mesh
protected the slope against possible local failures related to
drilling vibration. The bars were emplaced using central-
isers and grouted using a water cement ratio of 0.5 and a
grout pressure of 10 bars (1 MPa).
Studying the soil nailing pull out resistance, Yin et al.
(2009) found that for a grouting pressure of 130 kPa, the
maximum average soil–nail interface pressure was around
170 kPa for 5 mm nail head displacement, for a confining
stress of 200 kPa and for a soil saturation of 50 %. The
laboratory tests were carried in an instrumented 0.6
wide 9 0.8 high 9 1.0 m long chamber able to test only
one horizontal nail (1.2 m long and 40 mm diameter)
introduced in a 100 mm diameter hole. The soil was a
compacted completely decomposed granite fill. Figure 7
shows Yin et al.’s (2009) results of average soil–nail
interface shear strength rs as a function of grouting pressure
gp, confining stress and nail head displacement where an
interpolation line is included, which can be expressed as:
rs ¼ 50 þ 0:5gp in kPa ð7ÞAccording to (Eq. 7), the value adopted in the analysis
of gp should be 300 kPa to result in rs = 200 kPa, which is
actually beyond the studied limits of Yin et al. (2009).
Further research is needed to study the effect of higher
grouting pressures on the soil–nail interface shear strength.
Seismic analysis
It has been found that during seismic loading, a failure
mechanism develops in the same form as in the static case,
but the sliding soil blocks would be approximately double
the size (Tufenkjian and Vucetic 2000; Hanna and Juran
2000; Gaessler 2007). Comparing Figs. 6d and 8, it is clear
that these much larger sliding blocks would extend the
position of the failure surface, reducing the length of nail
able to develop shear resistance and hence the global factor
of safety (see Table 3).
Stability results
The factors of safety considered in the stability analysis for
the temporary soil nailing project are presented in Table 4
and the results of the stability analyses in Table 3. It is
clear that while the sliding and bearing capacity factors of
safety are higher than the required limits, the global factors
of safety are closer to the limits imposed by the project.
The minimum factor of safety does not occur in the seismic
case, but for the construction stage with an 8 m high wall
and three rows of nails. This highlights the importance of
the bottom row of nails which can mobilise shear resistance
for a longer length than the other nails.
Fig. 5 Unstable 4 m excavation showing material sliding in a
triangular shape
Table 2 Design parameters for the four rows of nails
Parameter Values
Length of anchors Ls, m 8, 8, 6, 6
Inclination of anchors, degrees 25, 20, 20, 15
Perforation diameter D, mm 110
Spacing between anchors SH, Sv, m 1.5, 1.8
Allowable tension capacity of bars Ta, kN 180
Soil-nail interface shear strength rs, kPa 200
Coefficient of horizontal acceleration kh, g 0.15
Overburden q, kPa 10
Nailed wall inclination b, degrees 85
Table 1 Geotechnical parameters of Maicillo
Parameter Values
Specific gravity Gs 2.708
Particle size d10, d30, d50, d60, mm 0.04, 0.25, 0.7, 1.1
Uniformity and curvature coefficients Cu, Cc 27.5, 1.42
Permeability coefficient k, m/s 6.4 9 10-4
Dry unit weight cd, kN/m3 16.4
Humidity W, % 8.2
Void ratio e 0.62
Saturation Sr, % 36
Cohesion c, kPa 13a
Peak angle of friction /’max, degrees 41.3a
Dilation angle W, degrees 14.8, 5.4, 4.8a
a Obtained for a constant normal stress of 25, 50 and 100 kPa
S. A. Villalobos et al.
123
The calculated tension loads on the steel nails are
summarised in Table 5. The first two rows correspond to
32.8 m long nails and rows 3 and 4 to 32.6 m long nails.
No yielding of the steel nails is expected as the tension
values are lower than the allowable value of 180 kN
(Table 2).
The ultimate nail pull-out capacity Tu can be obtained
assuming a uniformly mobilised soil–nail interface shear
strength rs, and a mean perforation diameter Ds = aD,
where a is a parameter related to the type of injection (1.2
is for an Injection Global and Unique IGU).
Tu ¼ pDsLrs ¼ 1:2pDLrs ð8Þ
Using a factor of safety of 1.5, the estimated allowable
pull-out capacity for the 6 and 8 m long nails is 330 and
440 KN, respectively.
Observations after a seismic event
The 27 February 2010 Chile earthquake provided a unique
opportunity to rethink the soil nailing design. The soil
nailing project had finished just 25 days before the 8.8
magnitude earthquake occurred at 3.34 a.m. local time. The
30 km deep hypocentre was located at Lat 36.2908S, Long
73.2398W, some 100 km north of the project site (Barri-
entos 2010). Unfortunately, no instruments were installed
4 m
2 m 2 m
2 m2 m
6 m
8 m 8 m
q = 10 kPa q = 10 kPa
q = 10 kPa q = 10 kPa
groundwatertable
groundwatertable
groundwatertable
groundwatertable
(a) (b)
(c) (d)
swimming pool
swimming poolswimming pool
swimming pool
Fig. 6 Construction sequence design
Soil nailing in granite
123
on site. However, accelerations were recorded in the centre
of Concepcion and in San Pedro de la Paz (see Fig. 3).
Figure 9a shows that at Concepcion, 2.5 km from the pro-
ject site, the maximum horizontal acceleration was 0.39 g
above sand and silt deposits 120 m deep. In San Pedro de la
Paz, 5.5 km from the project site, the maximum horizontal
acceleration was 0.63 g above residual soils from weath-
ered metamorphic rocks (see Fig. 3). It is also worth noting
the long duration of the seismic event and the lower fre-
quency content in Fig. 9a compared with Fig. 9b.
Attention is also drawn to the fact that around Concepcion
the earthquake induced a number of landslides in hill slopes
composed of the same heavily weathered granite (Verdugo
et al. 2010). However, inspections indicated although neither
cracks nor any sign of failure were present in the soil nailed
wall or in the swimming pool, a 5 mm wide, 100 mm deep
crack was found next to the swimming pool and running
23 m parallel to the masonry wall at a distance of 5.6 m. It is
believed that at least the upper block of the failure mecha-
nism was activated, but the soil nailing did not show evi-
dence of damage let alone indication of failure.
Given that the geology and type of soil in Quinta Junge
is closer to that found in San Pedro de la Paz than that in
the Concepcion centre, it can be assumed that the maxi-
mum horizontal acceleration was closer to 0.6 g than 0.4 g.
For that reason, the same pseudo static stability analysis
was undertaken with this acceleration effect. Figure 10
shows that the global factors of safety FSG = 1.61 and 1.32
for kh = 0 and 0.15, as previously presented in Table 3,
correspond to cohesion and angle of friction of 10 kPa and
30�, respectively. However, from this analysis, a soil
nailing collapse should have occurred for kh C0.3, resulting
in FSG = 0.75 for kh = 0.6. In view of the fact that failure
did not occur, it seems that the use of an undrained angle of
friction value of 30� resulted in a conservative design. The
other two curves in Fig. 10 correspond to drained values
c = 13 kPa and / = 37.5 and 41.3�, which are more likely
to have been mobilised during the earthquake since the soil
had a moisture content of\8 % (February being the driest
month of the year). These latter values were measured in
direct shear tests (see Fig. 4c; Table 1). The use of these
drained values led to global factors of safety above 1 for
accelerations near to 0.6 g.
The lower curve in Fig. 10 represents the worst scenario
when the soil is saturated and large accelerations occur
Fig. 7 Average interface shear strength versus grouting pressure
(data from Yin et al. 2009)
q = 10 kPa
groundwatertable
8 m
2 m swimming pool
Fig. 8 Seismic design using Kh = 0.15 for temporary soil nailing
Table 3 Tempory factors of safety calculated
Case Constr.
1
Constr.
2
Constr.
3
Static Seismic
Sliding 33.1 16.8 10.5 10.5 7.8
Bearing
capacity
9.6 4.6 4.2 4.2 4.0
Global 1.44 1.33 1.28 1.61 1.32
Table 4 Factors of safety limits for the tempory soil nailing project
Case Construction Static Seismic
Sliding 1.3 1.3 1.1
Bearing capacity 2.5 2.5 2.3
Global 1.2 1.35 1.1
Table 5 Calculated tension loads in steel nails in kN
Nails row Constr. 1 Constr. 2 Constr. 3 Static Seismic
1 125 135 140 135 175
2 – 135 140 115 135
3 – – 140 115 135
4 – – – 115 140
S. A. Villalobos et al.
123
during 1 or 2 min in a strong earthquake. This situation
might occur if an earthquake takes place during the rainy
season.
Conclusions
This paper presents the design and construction of a soil
nailing project in a heavily weathered granite. The stability
analysis was based on a limit equilibrium method and the
results interpreted in terms of tension loads in the nails and
factors of safety. The failure mechanism consists of sliding
blocks which develop shear strength on the contact areas
with a geometry based on experimental findings. Nails
assist in increasing the shear resistance and hence the
stability of the excavated slope. It was found that at the final
stage of construction, when the bottom row of nails was not
yet installed, the situation is even less favourable than the
seismic case with kh = 0.15 for the finished soil nailing.
The analysis, based on calculations of resistant and
applied loads, proved adequate for the design of nailed
walls under drained conditions, even under a high seismic
loading. However, it is important to consider deformations
and displacements experienced by the soil, nails and wall,
during construction, service and under seismic loading.
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Soil nailing in granite
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