Site response and liquefaction hazard analysis of Hawassa town, Main Ethiopian Rift OPEN ACCESS EDITED BY

Abstract

The study area is located in one of the most earthquake prone regions in southern Ethiopia, which is characterized by small-to-intermediate earthquake occurrences causing damage to buildings. Predicting liquefaction hazard potential and local site effects are imperative to manage earthquake hazard and reduce the damage to buildings and loss of lives. The objectives of this work were to perform the equivalent linear response analysis (ELA) and shear wave velocity (Vs.)-based liquefaction hazard analysis and classify the site into different seismic site classes based on the European and American codes. The SPT-N and Vs.30 values showed the site falls in the C and D classes based on the NEHRP (2015) code but falls in the B and C classes based on the EC8 (2003) code. The susceptibility of liquefaction was evaluated using grain size analysis curves. Moreover, peak ground acceleration (PGA), spectral acceleration (SA), and maximum strain (%), which are very critical to understanding the local site effects, were estimated by the DeepsoilV.7 program. The cyclic stress ratio and cyclic resistance ratio were used to calculate the factor of safety (FS). A liquefaction potential index (LPI), probability of liquefaction (PL), and probability of liquefaction induced ground failure (PG) were used to assess the probability of liquefaction. The peak ground acceleration (g) values ranged from 0.166 to 0.281 g, whereas spectral acceleration (g) was found to be high at 0.1–1s. The liquefaction susceptibility screening criteria revealed that the study area is highly susceptible to liquefaction. FS is < 1 for a liquefied site, but FS is > 1 for non-liquefied sites. In comparison to non-liquefied sites, the liquefaction forecast site has a liquefaction potential index value of 0–54.16, very likely high PL, and high PG. The findings will be helpful in the design of structures and in solving practical challenges in earthquake engineering.

Full text

Site response and liquefaction
hazard analysis of Hawassa town,
Main Ethiopian Rift
Alemayehu Ayele
1
,
2
*, Matebie Meten
1
and Kifle Woldearegay
3
1
Department of Geology, College of Applied Science, Addis Ababa Science and Technology University,
Addis Ababa, Ethiopia,
2
Department of Geology, College of Natural and Computational Science,
Wachamo University, Hosanna, Ethiopia,
3
School of Earth Sciences, Mekelle University, Mekelle,
Ethiopia
The study area is located in one of the most earthquake prone regions in
southern Ethiopia, which is characterized by small-to-intermediate earthquake
occurrences causing damage to buildings. Predicting liquefaction hazard
potential and local site effects are imperative to manage earthquake hazard
and reduce the damage to buildings and loss of lives. The objectives of this work
were to perform the equivalent linear response analysis (ELA) and shear wave
velocity (Vs.)-based liquefaction hazard analysis and classify the site into
different seismic site classes based on the European and American codes.
The SPT-N and Vs.30 values showed the site falls in the C and D classes based
on the NEHRP (2015) code but falls in the B and C classes based on the EC8
(2003) code. The susceptibility of liquefaction was evaluated using grain size
analysis curves. Moreover, peak ground acceleration (PGA), spectral
acceleration (SA), and maximum strain (%), which are very critical to
understanding the local site effects, were estimated by the
DeepsoilV.7 program. The cyclic stress ratio and cyclic resistance ratio were
used to calculate the factor of safety (FS). A liquefaction potential index (LPI),
probability of liquefaction (PL), and probability of liquefaction induced ground
failure (PG) were used to assess the probability of liquefaction. The peak ground
acceleration (g) values ranged from 0.166 to 0.281 g, whereas spectral
acceleration (g) was found to be high at 0.1–1s. The liquefaction
susceptibility screening criteria revealed that the study area is highly
susceptible to liquefaction. FS is <1 for a liquefied site, but FS is >1 for non-
OPEN ACCESS
EDITED BY
Mahmood Ahmad,
University of Engineering and
Technology, Peshawar, Pakistan
REVIEWED BY
Denise-Penelope N. Kontoni,
University of the Peloponnese, Greece
Beenish Jehan Khan,
CECOS University of Information
Technology and Emerging Sciences,
Pakistan
*CORRESPONDENCE
Alemayehu Ayele,
alex98geo@gmail.com
SPECIALTY SECTION
This article was submitted to Earthquake
Engineering,
a section of the journal
Frontiers in Built Environment
RECEIVED 08 August 2022
ACCEPTED 31 August 2022
PUBLISHED 26 September 2022
CITATION
Ayele A, Meten M and Woldearegay K
(2022), Site response and liquefaction
hazard analysis of Hawassa town, Main
Ethiopian Rift.
Front. Built Environ. 8:1014214.
doi: 10.3389/fbuil.2022.1014214
COPYRIGHT
© 2022 Ayele, Meten and Woldearegay.
This is an open-access article
distributed under the terms of the
Creative Commons Attribution License
(CC BY). The use, distribution or
reproduction in other forums is
permitted, provided the original
author(s) and the copyright owner(s) are
credited and that the original
publication in this journal is cited, in
accordance with accepted academic
practice. No use, distribution or
reproduction is permitted which does
not comply with these terms. Abbreviations: CSR, Cyclic Resistance Ratio; CSR, Cyclic Stress Ratio; EAR, East African Rift System;
EC8, Eurocode-8; ELA, Equivalent Linear Response analysis; FC, Fines content; FD, Frequency
Domain; FS, Factor of Safety; GWT, Groundwater table; LPI, Liquefaction Potential Index; MER,
Main Ethiopian Rift System; Mw, moment magnitude; NEHRP, National Earthquake Hazards
Reduction Program; PEER, Pacific Earthquake Engineering Research Center; PG, Probability of
liquefaction induced ground failure; PGA, Peak Ground Acceleration; PL, Probability of
Liquefaction; SA, Spectral Acceleration; SMER, Southern Main Ethiopia Rift; SPT-N, Standard
Penetration test blow counts; TWBH1- 3 m, Tikuri woha Borehole 1 grain size curve at 3 m depth;
TWBH1-6 m, Tikuri woha Borehole 1 grain size curve at 6 m depth; TWBH1-13 m, Tikuri woha
Borehole 1 grain size curve at 13 m depth; TWBH2-3 m, Tikuri woha Borehole two grain size
curve at 3 m depth; TWBH2-6 m, Tikuri woha Borehole two grain size curve at 6 m depth;
TWBH2-11 m, Tikuri woha Borehole two grain size curve at 11 m depth; ADBH1-3 m, Adare
Borehole 1 grain size curve at 3 m depth; ADBH1-11 m, Adare Borehole 1 grain size curve at 11 m
depth; ADBH2-3 m, Adare Borehole two grain size curve at 3 m depth; ADBH2-13 m, Adare Borehole
two grain size curve at 13 m depth; Vs, Shear wave velocity; Vs.30, Average shear wave velocity up to a
depth of 30 m.
Frontiers in Built Environment frontiersin.org01
TYPE Original Research
PUBLISHED 26 September 2022
DOI 10.3389/fbuil.2022.1014214

liquefied sites. In comparison to non-liquefied sites, the liquefaction forecast
site has a liquefaction potential index value of 0–54.16, very likely high PL, and
high PG. The findings will be helpful in the design of structures and in solving
practical challenges in earthquake engineering.
KEYWORDS
shear wave velocity, factor of safety, equivalent linear response analysis, liquefaction
susceptibility, liquefaction hazard evaluation
Introduction
Globally, earthquake hazard is causing thousands of victims
and deaths, hundreds of billion dollars of infrastructure damages,
and environmental losses (Pirhadi et al., 2018;Irinyemi et al.,
2022). The earthquake activities in Ethiopia are mostly associated
with the Afar depression, the escarpment and the Ethiopian rift
system as these areas are seismic source zones due to their
tectonic, geologic and seismic characteristics (Gouin, 1979;
Mammo, 2005;Ayele, 2017;Wilks et al., 2017;Fentahun
et al., 2021). The main Ethiopian rift (MER) system is mainly
responsible for triggering earthquakes in Ethiopia (Wilks et al.,
2017;Lamessa et al., 2019). It has resulted in the loss of human
lives and damage to infrastructure (Gouin, 1979;Kebede and Van
Eck, 1997;Mammo, 2005;Ayele et al., 2021).
Liquefaction is one of the most devastating seismic hazards
related to earthquakes in the world. It occurs when earthquake
shaking and increased pore water pressure reduce the strength
and stiffness of soil (Marcuson, 1978;Papathanassiou et al., 2011;
Tehran et al., 2016;Bahari et al., 2020;Ansari et al., 2022).
Therefore, the evaluation of the liquefaction susceptibility and its
resistance is an important component of seismic hazard
assessment in an earthquake-prone region. In many towns in
Ethiopia, including the research area, a recent earthquake with a
moment magnitude (Mw >5) in 2016 of the Gregorian calendar
(G.C) caused damage to buildings (Wilks et al., 2017;Lamessa
et al., 2019) but no research in the study area has yet been done
on the prediction of earthquake-induced liquefaction hazard and
local site effects analysis, which is aimed at minimizing such
earthquake hazards. Earthquake-induced liquefaction can cause
the failure of foundations, soil embankments, and dams,
especially in cities built on young alluvial and lacustrine
sediments. These failures ultimately affect the social and
financial status of the region (Pokhrel et al., 2013;Ahmad
et al., 2019;Subedi and Acharya, 2022). As a result, the
damage from earthquake-induced liquefaction is often worse
than the damage caused by the other effects of an earthquake
around the world. Therefore, sites that may be highly prone to
earthquake-induced liquefaction and local site effects related to
amplification should be identified in order to reduce the damage
from the earthquake disaster and loss of lives. The liquefaction
potential of soils depends on grain size distribution, fines content,
geological time, sedimentation, permeability, earthquake
magnitude, and earthquake duration (Özaydın, 2007;Meisina
et al., 2022). Therefore, identifying the liquefaction susceptibility
criteria is very important to evaluate the area with high
susceptibility to liquefaction.
The liquefaction potential evaluation of soils is determined
using laboratory (Polito, 2001;Bray et al., 2004;Li et al., 2022)
and in-situ tests (Rahman and Siddiqua, 2017;Subedi and
Acharya, 2022). Due to its high cost and the difficulty of
bringing undisturbed samples and conducting high quality
tests, the simplified procedure based on in-situ tests such as
standard penetration tests (SPT), cone penetration tests (CPT),
and Vs. measurements is widely practiced in the United States
and most parts of the world. The simplified procedure, which
was originated by Seed and Idriss (1971),usesSPTdatafor
evaluating the liquefaction potential of granular soils (Seed
et al., 1984;Youd et al., 2001;Youd and Idriss, 2001;Rahman,
2019). Over the years, the simplified procedure has been revised
and updated with additional data and has become the most
commonly used way to assess the liquefaction potential of
granular soils (Harder, 1997;Robertson and Wride, 1998;
Goren and Gelisli, 2017;Bahari et al., 2020). The CPT is
used for liquefaction hazard evaluations (Robertson and
Wride, 1988;Ahmad et al., 2021), which has also been
revised and updated by Seed and de Alba (1986),Stark and
Olson (1995),Olsen (1997),Robertson and Wride (1998),and
Boulanger and Idriss (2014). The liquefaction potential
evaluation based on the SPT and CPT data are fairly well
developed. However, the penetration tests may be
impractical or unreliable at some sites when penetration
tests are not conveniently performed at all depths. A
promising alternative to the deterministic-based approaches
is provided by in situ measurements of small-strain Vs. to
estimate the liquefaction potential of soils (Andrus and Stokoe,
1997;Andrus and Stokoe, 2000;Zhang, 2010;Ji et al., 2021).
This study applied Vs. measurements using Multichannel
analysis of surface wave (MASW) as; 1) the Vs.
measurements are possible in hard soils where the SPT and
CPT are difficult to penetrate or to collect undisturbed stiff soil
samples or at the site where these may not be permitted; 2) Vs. is
a basic mechanical property of soil materials directly related to
the small-strain shear modulus; 3) the small-scale shear
modulus is a parameter required for estimating the dynamic
properties of soil and soil structure interaction analyses (Youd
et al., 2001); and 4) it is non-invasive, cost-effective, and covers
large areas in a short period of time. A factor of safety (FS) is
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used to determine the liquefaction potential (Uyanıketal.,
2013;Ji et al., 2021). Iwasaki, (1978) has suggested a
liquefaction potential index (LPI) to predict the severity of
liquefaction at a specific site. A LPI is estimated by a FS against
liquefaction at different depths. Equivalent linear response
analysis (ELA) has been applied to determine peak ground
acceleration (PGA) to estimate the cyclic stress ratio (Satyam
and Towhata, 2016;Putti and Satyam, 2018;Hashash et al.,
2020). The study area is located in the basin of thick lacustrine
sediments, which can amplify the earthquake ground motion
and may probably cause an earthquake-induced liquefaction
hazard. Its vicinity to active seismic zones, rapid
industrialization, poor control of construction practices and
quality, lack of site effect information, and increased population
growth rate make the study area highly susceptible to
earthquake damage in the future. Therefore, this demands
ELA and liquefaction potential analysis in Hawassa town. In
this study, we used Vs-based liquefaction potential evaluation
and ELA by supplementing SPT values in Hawassa town. In
addition,wealsousedFSagainstliquefaction,LPI,probability
of liquefaction (PL), and probability of ground failure (PG) to
estimate the probability of liquefaction in the study area. The
objectives of this work are to perform the equivalent linear
response analysis (ELA), evaluation of liquefaction
susceptibility, Vs-based liquefaction hazard analysis, and
classify the site into different engineering classes based on
the European and American codes. The Vs. and SPT were
used to characterize the study area, whereas ELA was applied to
determine PGA (g), 5% damped SA (g) and maximum strain
(%) using the DeepsoilV.7 program. FS was utilized to evaluate
the susceptibility of liquefaction. In addition, LPI, PL, and PG
were used in this research work to evaluate the probability of
liquefaction. The results showed amplification, probability of
liquefaction, and ground deformation occuring at a shallow
depth.
Materials and methodology
Location
Hawassa town is located in the southern branch of the Main
Ethiopian Rift (MER) system. The study area is geographically
located at longitude 7°1′to 7°5′N and latitude 38°28′to 38°29′E.
The MER system is a part of the East African Rift (EAR).
Furthermore, Hawassa town is situated on the eastern shore
of Lake Hawassa (Figure 1).
Seismotectonics
The MER forms an active plate boundary between the Africa
(Nubia) and Somalia plates in the northern East African Rift
(EAR) system. In addition, the MER is an example of mature
FIGURE 1
Location map of the study area.
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continental rifting at the northernmost part of the EAR. Thus, the
EAR is one of the most geologically active features on Earth
(Chorowicz, 2005;Lamessa et al., 2019). The MER is divided into
northern, central and southern main sectors (Agostini et al.,
2011;Wilks et al., 2017) based on the lithospheric characteristics
(Keranen et al., 2004), crustal thickness and modification
(Keranen and Klemperer, 2008). Hawassa town is located in
the southern main Ethiopia rift (SMER) segment, which is largely
fault controlled and actively deformed (Corti et al., 2013). As a
result, it has faced low to intermediate magnitude earthquakes
(Mw >5) in recent years, which caused damage to buildings and
created public concern in Hawassa town (Wilks et al., 2017).
Geology
Hawassa geology is composed of upper Miocene to
Pliocene volcanic rocks of the Nazret group dominated
by rhyolitic ignimbrites that form the basement of the MER
(Woldegabriel, 1990;Žáček et al., 2014). According to
Žáček et al. (2014), the lithological units in Hawassa
town and its environs include (Figure 2)fluvial
sediments, alluvial sediments, polygenetic sediments, Wondo
Koshepumicefallandflow deposits, scoria cones, tuff cones,
Hawassa basalts, and Hawassa rhyolitic ignimbrites. A total of
19 Multichannel Analysis of Surface Wave (MASW), 10 SPT
data (ARCON Design Build plc, 2018), and 10 BH (borehole)
data (SDCSE, 2019) points whichwereusedinthisstudyare
shownonthegeologicalmap(Figure 2).
Methodology
The Vs. is the geotechnical parameter used to measure
the mechanical properties of a soil. The Vs.30 is the travel
time averaged shear wave velocity in the topmost 30 m of
depth. In this study, insitu Vs. measurements were conducted
at the different sites of Hawassa town using MASW in the
field. MASW is a surface wave analysis method which is used
to measure the Vs. at a surveyed site (Kramer, 1996;Nath and
Jakka, 2012;Putti and Satyam, 2018;Kamel and Badreddine,
2020). Then, Vs.30 was calculated using Eq. 1 (Boore, 2004):
Vs.30 30 m
N
i
hi
Vi
,(1)
where Vs.30 is the shear wave velocity up to a depth of 30 m, Vi is
the shear wave velocity for the ith layer in m/s, hi is the thickness
of the ith soil layer in meters, and N is the number of layers in the
top 30 m soil strata. SPT-N values are used to evaluate the
dynamic properties of soils. The SPT was conducted inside
boreholes using a standard hammer by weighing 63.5 kg and
FIGURE 2
Geological map of the study area and its environs (modified after Žáček et al., 2014) with MASW, SPT, and BH data points.
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falling freely from a height of 760 mm based on the procedure of
British standards (Figure 3A). Figure 3B shows SPT-N values of
the last 300 mm of penetration for the corresponding depths in
the wooden core boxes. A total of 19 MASW data points were
collected for site characterization and classification (Figure 2).
However, five typical sites of MASW data points were chosen for
the detailed ELA and liquefaction hazard analysis by
supplementing 10 SPT and 10 borehole (BH) data.
Deterministic liquefaction hazard
analysis
Liquefaction susceptibility
The liquefaction susceptibility is evaluated based on the
properties of the soil. Many researchers (Tsuchida, 1970;
Kramer, 1996;Seed et al., 2003;Bahari et al., 2020) have used
FIGURE 3
(A) Field investigations and excavation (B) SPT samples and wooden core boxes.
FIGURE 4
Limits in the gradation curves separating liquefiable and non-liquefiable soils (after Tsuchida, 1970).
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the particle size, shape and gradation of soil to evaluate the
susceptibility of liquefaction, while some research has reported
the age of the deposit and depth of the groundwater level for the
evaluation of susceptibility to liquefaction (Youd and Hoose,
1977;Youd and Perkins, 1987;Kavazanjian et al., 1997;Satyam
and Towhata, 2016;Naik et al., 2020;Subedi and Acharya, 2022).
In this study, the site’s susceptibility to liquefaction was evaluated
according to (Tsuchida, 1970), which is very easy, acceptable
worldwide, and shows gradation curve boundaries for most
liquefiable and non-liquefiable soils (Figure 4).
FS
A FS is used to evaluate liquefaction hazards using Vs. Several
studies were used to estimate FS based on the Vs. data (Andrus and
Stokoe, 2000; Oshnavie and Khalkha, 2019; Bahri et al., 2020;Ji et al.,
2021)anditwascalculatedusingEq. 2 (Seed and Idriss, 1982).
FS CRR
CSR,(2)
where FS is the factor of safety, CSR is the cyclic stress ratio, and
CRR is the cyclic resistance ratio. When the FS is >1, it is
assumed that no liquefaction will occur, but when the FS is <1,
liquefaction will occur. The CSR is the average cyclic shear stress
(τav) of soils due to cyclic or earthquake loading to the vertical
effective stress (σv.) acting on a soil layer (Seed and Idriss, 1971;
Andrus and Stokoe, 2000;Liu et al., 2001;Andrus et al., 2003;
Uyanık et al., 2013;Rahman, 2019;Ji et al., 2021) and thus it was
estimated using Eq. 3 (Seed and Idriss, 1971) at any depth.
CSR τav
σv.0.65PGA max
gσv
σv.rd, (3)
where τav is the average cyclic shear stress which is caused by
the earthquake, and it is assumed to be 0.65 of the maximum
induced stress, PGA max is the peak horizontal acceleration at
thegroundsurfaceandisdeterminedfromsiteresponse
analysis using ELA, g is the acceleration of gravity (9.81 m/
s
2
), σv. is the effective vertical stress at a depth, σvis the vertical
stress at a depth, and rd is the depth reduction factor. The stress
reduction coefficient (rd) is a function of site stratigraphy, soil
properties, and the characteristics of motion excitation where
‘‘rd’’ is applied to adjust the flexibility of soil (Seed and Idriss,
1971; Idriss and Bounger, 2004; Grasso et al., 2020). Because it
is a simple, widely accepted, and linearly decreasing function to
a depth, the rd was calculated using Eq. 4 (Iwasak, 1986).
rd 1−0.015z, (4)
where rd is stress reduction coefficient and z is a depth.
CRR, which is the cyclic resistance of a soil during an earthquake,
is determined using the relationship between measured Vs., stress-
corrected shear wave velocity (Vs1) and magnitude scaling factor
(MSF) (Andrus and Stokoe, 1997;Youd et al., 2001;Uyanık, 2002;
Uyanık and Taktak, 2009;Ortiz-Hernández et al., 2022)andthusCRR
calculated using Eq. 5 (Andrus and Stokoe 2000).
CRR 0.022Vs1
1002
+2.81
Vs1
*−Vs1−1
Vs1
*MSF, (5)
where Vs1
*the limiting upper value for cyclic liquefaction
occurrence, MSF is the magnitude scaling factor to account
for the effect of earthquake magnitude. Vs1
*is related to
average fines content (FC) and expressed by Eqs 6a,6b,6c.
Vs1
*250 m
sf or sands with FC ≤5%,(6a)
Vs1
*215 −0.5(FC −5)ms,f or sands with 5 <FC <35%,
(6b)
Vs1
*200 m
s,f or sands and silts with FC ≥35%,(6c)
where FC is the average fines content in percent by mass. The
gradation curve showed the FC >40%. As a result, Eq. 6c was used
to estimate theVs1
*.
To estimate Vs1 of soils (Sykora, 1987;Robertson et al., 1992;
Uyanık et al., 2013) have used measured Vs. Thus, Vs.1 was
calculated using Eq. 7 (Andrus and Stokoe, 2000).
Vs1Vspa
σv.0.25
,(7)
where Vs1 the stress-corrected shear wave velocity (m/s), Vs. is
the measured shear wave velocity (m/s), Pa is 100 kN/m
2
reference stress or atmospheric pressure, and σv. is effective
vertical stress (kN/m
2
). The MSF is used to adjust CRR for an
earthquake Mw greater or less than 7.5. Therefore, Eq. 8 (Andrus
and Stokoe, 1997;Andrus and Stokoe, 2000) was utilized to
calculate MSF.
MSF Mw
7.5n
,(8)
where MSF is the magnitude scaling factor, Mw is the moment
magnitude, and n is the standard, which depends on the
amount of earthquake Mw. The standard value of n
is −3.3 for Mw <7.5 and n is −2.5 for Mw >7.5 (Andrus
and Stokoe, 1997;Andrus and Stokoe, 2000). Wilks et al.
(2017) and Ayele (2017) have suggested the earthquake Mw of
the study area has occurred in the past was less than 7.5.
Therefore, the standard value of −3.3 was used. For the
estimation of liquefaction potential index (LPI), Iwasaki
et al. (1982) proposed equation for LPI, which was later
summarized by Iwasaki, (1986),wasused(Eqs 9a,9b,9c,
9d,and9e).
LPI 20
0
F(z).W(z)dz,(9a)
F(z)1−FS f or FS <1,(9b)
F(z)0 f or FS ≥1,(9c)
W(z)10 −0.5z for Z <20 m, (9d)
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W(z)0 f or Z >20 m, (9e)
where z is the depth of the midpoint of the soil layer, dz is the
differential increment of depth, w (z) is the weighting factor, and
F (z) is the severity factor. Eq. 10 shows the probability of
liquefaction (PL) at each site was calculated using Eq. 10
(Juang et al., 2003).
PL 1
1+FS
0.964.5,(10)
where PL is the probability of liquefaction and FS is the factor of
safety. PG is required information for making risk-based design
decisions. Therefore, the PG for near foundation was estimated
using Eq. 12 (Li et al., 2006).
PG 1
(1+e4.9−0.74LPI),(12)
where PG is probability of liquefaction induced ground failure
and LPI is liquefaction potential index.
ELA
ELA was used to model non-linear soil parameters in
terms of equivalent linear soil properties. The earlier
researchers used peak ground acceleration (PGA),
spectral acceleration (SA), and maximum strain (%) to
predict the ground motion (Kramer, 1996;Yoshida et al.,
2001;Eker et al., 2015;Putti and Satyam, 2018;Hashash
et al., 2020;Soltani-Azar, 2022). Due to the unavailability
of strong ground motion in Ethiopia, the Northridge
ground motion (Mw 6.69) was consulted as an input
motion from the Pacific Earthquake Engineering
Research Center (PEER, 2010) database (http://peer.
berkeley.edu) based on the target response spectrum of
the region (Kramer, 1996;Kebede and Van Eck, 1997;
Mammo, 2005;Yee et al., 2013;Alemu et al., 2018;
Bahari et al., 2020). In addition, several researchers
adopted the material properties from the families of
normalized shear modulus reduction and damping
curves of the DEEPSOILV.7 program for cohesionless
soil (Seed and Idriss, 1970;Seed et al., 1986;Idriss,
1990;Mammo, 2005;Carlton and Abrahamson, 2014;
Alemu et al., 2018;Rahman, 2019;Nguyen et al., 2020).
Likewise, the families of normalized shear modulus
reduction and damping curves developed by Seed et al.
(1986) were consulted for this study, which were later
modified (Darendeli, 2001)(Figures 5A,B).
Results and discussion
Site characterization and classification
A total of 19 MASW were conducted for the site
characterization and classification to evaluate the local site
conditions. The Vs.30 was used to characterize and classify
the research region according to the seismic design codes. The
estimated Vs.30 values of the study area range from 248.9 m/s to
371.3 m/s as shown in Table 1. The site class was classified as C
and D classes based on the NEHRP (2015) code, but B and C
classes were classified based on the EC8 (2003) code. The MASW
data showed that 17 surveyed sites belong to the D class for
NEHRP (2015) but C based on the EC8 (2003), whereas two sites
were found to be C according to NEHRP (2015) (Figure 6A) and
B based on the EC8 (2003) (Figure 6B). In addition, the study
area was dominated by D seismic classes for NEHRP (2015) and
C seismic classes for EC8 (2003). As compared to the previous
studies (Eurocode-8, 2003;BSSC, 2015;Alemu et al., 2018), this
study was given similar seismic site classes.
The results in Figure 7 showed the gradation curve in the
research area. It was plotted that the soil samples taken from the
study area were at a depth of 3 m, 6m, and 12 m. As it was seen
from the results, the plotted grain size analysis curves of Hawassa
town showed silty sand soil. From the graphs, the percentage of
fines content (FC) was greater than 40%. The findings of this
work are similar to those of Alemu et al. (2018).
The SPT-N values at different depths and the lithological
descriptions are shown at Table 2. Within the investigated depth
of 0–13 m, volcanic materials such as silty sand soil profile were
FIGURE 5
(A) Normalized modulus reduction (G/Gmax) and (B) material
damping (D) curve (Seed et al., 1986) and later modified by
(Darendeli, 2001).
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encountered. The volcanic material demonstrated variegated, light
brown, and dull color. The SPT-N values revealed the resedimented
pryroclastic material from loose/soft to medium stiff.
Table 3 showed the laboratory results of soil in the study area.
Each soil sample in the study area was classified based on the
unified soil classification (USC). The Atterberg limit value was
not possible (NP) to determine for each soil sample due to the
absence of the plasticity property of soil. Thus, each soil sample in
the research area showed silty sand (SM) soil. The FC (%) was
greater than 40% for the analyzed soil in the study area. The
results of this study are also the same as the previous results of
Alemu et al. (2018).
ELA
Ground motion prediction parameters were determined
using PGA (g) and SA (g) at the typical five sites. The
Northridge ground motion with Mw 6.69 from the PEER
database was used as an input to simulate the propagation of
TABLE 1 The MASW data of the study area.
S.No Site code Vs.30 (m/s) Study area site classes
NEHRP (2015) EC 8 (2003)
1 1 362.3 C B
2 2 350 D C
3 3 301.1 D C
4 4 269.3 D C
5 5 371.3 C B
6 6 301.1 D C
7 7 345.5 D C
8 8 265.8 D C
9 9 260.7 D C
10 10 267.9 D C
11 11 270.5 D C
12 12 275.7 D C
13 13 355.9 D C
14 14 356.9 D C
15 15 270.4 D C
16 16 347.9 D C
17 17 266.9 D C
18 18 264.7 D C
19 19 248.9 D C
FIGURE 6
(A) Histogram of NEHRP (2015) site classification and (B) Histogram of site EC8 (2003) classification.
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FIGURE 7
Grain size curves of the soil in the study area.
TABLE 2 SPT-N values and lithological descriptions.
Borehole ID Depth (m) SPT-N values Lithological description
TWBH1 2–6 7 Variegated color, loose, non-plastic silty sand soil
6–13 10 Light brown color, soft to medium stiff, non-plastic silty sand soil (resedimented pyroclastic material)
TWBH2 3–6 8 Variegated color, loose, non-plastic silty sand soil
6–11 11 Dull light, medium dense, non-plastic silty sand soil (volcanic ash)
ADBH1 2–4.45 7 Variegated color, soft to medium stiff, non-plastic silty sand soil layer
4.45–11 10
ADBH2 2–4.45 11 Variegated color, medium stiff, non-plastic silty sand soil
4.45–13 12
TABLE 3 Summarized geotechnical results in the study area.
Borehole ID Depth (m) Atterberg Limit % Soil
classification
(USCS)
LL PL PI Gravel Sand Fine
BH-1Adr 3 NP NP NP - 60 40 silty sands SM
TWBH2 3 NP NP NP - 36 64 silty sands
ADBH2 3 NP NP NP - 59 41 silty sands
TBH1 6 NP NP NP - 23 77 silty sands
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seismic waves from engineering bedrock to the ground surface.
The soil thickness, groundwater table, density, and Vs. were used
to simulate earthquake ground motion at a depth of 30 m since it
is used as an input to evaluate local effects at a site. Then, for all
five typical sites, the variation of PGA (g), maximum strain (%)
and SA (g) at different depths were generated. At the ground
surface, the generated PGA (g) values in Figure 8A showed: (i)
0.281 g for site 1, (ii) 0.234 g for site 2, (iii) 0.203 g for site 3, (iv)
0.166 g for site 4 and (v) 0.18 g for site 5. As a result, all the sites
increased the reference input motion due to the local site effects.
The results showed that the higher PGA (g) value was observed at
site 1 as compared to other sites due to the shallow occurrence of
the groundwater table (Rahman and Siddiqua, 2017;Alemu et al.,
2018) The small increases of PGA (g) values were observed at site
4. In addition, the PGA (g) for site 1, site 2, site 3, site 4 and site
5 showed very low values at a depth range between 5 and 11 m.
However, PGA (g) was found higher from the ground surface to a
depth of 5 m for the five sites (Figure 8A). The plots of depth
versus PGA (g) for the five sites followed similar patterns. The
typical five sites in the study area showed high amplification from
the considered depth to the surface for the input motion.
The earthquake ground motion was greatly influenced by site
local soil conditions, as shown by PGA values in Figure 8A. This
result is also comparable to Alemu et al. (2018) but it showed
more PGA (g) value as compared to other research works
(Mammo, 2005;Ayele, 2017). As a result, the PGA (g) values
for Hawassa town have the potential to amplify an earthquake
ground motion to a significant degree, even more than the one
provided by the seismic codes of Ethiopia (ES EN 1998:2015),
which may aggravate the occurrence of earthquake-induced
liquefaction. This result implies that the study area has got a
higher PGA (g) value than the one specified by the well seismic
standards (ES EN 1998:2015; Alemu et al., 2018;Puti and Satyam,
2018).
The maximum strain (%) for the five sites of Hawassa town
was indicated (Figure 8B) at site 1; the maximum strain (%) was
found to be increased at a depth of 5 m and then decreased up to
the depth of investigation; at site 2, the maximum strain (%) was
increased at a depth of 8 m and then showed a decreasing trend at
a depth of 22 m before increase; at site 3, the maximum strain (%)
was increased at a depth of 9 m, decreased at a depth of 23 m and
finally started to increase at a depth of 28 m; at site 4, the
maximum strain (%) was increased up to depth of 18 m
before starts to a decreasing and at site 5, the maximum
strain (%) was increased up to depth of 10 m and then it was
decreased from a depth of 10–30 m. In addition, a higher strain
was observed at site 1 than other sites whereas low strain was
exhibited at site 5 (Figure 8B). Therefore, larger stress
deformation will be expected for site 1 than the other sites for
the simulated earthquake ground motion. The 5% damped SA (g)
for the five sites of Hawassa town showed very high values at a
period of 0.1–1s(Figure 8C). At the period of 0.1–1 s, the 5%
damped SA (g) of site 1, site 2, site 3, site 4 and site 5 was
forecasted to be very high for the input ground motion. The
analysis of 5% damped SA (g) was found to be 1.2 g for site 1,
0.81 g for site 2, 0.8 g for site 3, 0.71 g for site 4 and 0.7 g for site 5.
The higher value of 5% damped SA (g) was observed at site
1 whilst the lower value of 5% damped SA (g) was shown at site 5.
The highest value of 5% damped SA (g) at site 1 indicates that
there will be high amplification for the earthquake Mw 6.69.
From Figures 8A–C, it was also observed that the 5% damped SA
(g) at the typical sites was nearly similar to that of Alemu et al.
(2018).
Liquefaction susceptibility criteria
The soil samples were analyzed to determine the
susceptibility of liquefaction soil in the study area. Thus, the
geotechnical characteristics of different soil samples were
determined using the grain size distribution curves. The curves
indicated the grain size of soils with engineering properties
(Figure 9). In addition, the grain size curves also showed that
the soils in all the analyzed samples were silty sand soils. As a
FIGURE 8
(A) PGA (g), (B) maximum strain (%) and (C) 5% damped SA (g)
for site 1, site 2, site 3, site 4 and site 5.
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result, the curve indicated the high liquefaction susceptibility of
these soils (Tsuchida, 1970;Rahman and Siddiqua, 2017) as per
the screening criteria. The grain size curves demonstrated all the
soil samples in the research area at a depth of 1–13 m (Figure 9)
fall at the boundaries of most liquefiable soil (Tsuchida, 1970)
according to the preliminary screening liquefaction susceptibility
criteria. In addition, the grain size analysis of the silty sand soil
demonstrated that the soils are susceptible to liquefaction as per
the screening criteria (Tsuchida, 1970;Rahman and Siddiqua,
2017;Alemu et al., 2018). As a result, the liquefaction
susceptibility of soils mainly depends on the soil type of the
site. The grain size distribution curve in this study exhibited high
liquefaction susceptibility at the boundaries of most liquefiable
soils (Figure 9) as compared to earlier research works (Seged and
Haile, 2010;Satyam and Towhata, 2016).
The study area comprised silty sand soil with a relative density of
less than 50% based on the geotechnical data (ARCON Design Build
plc, 2018). Therefore, liquefaction susceptibility was expected in the
silty sand soil according to the screening criteria (Kramer, 1996;
Kavazanjian et al., 1997). This is because the soil with a lower relative
density was more susceptible to liquefaction than that with a soil
higher relative density. The depth of groundwater in the study area
ranges from 2 to 56 m (SDCSE, 2019). Therefore, the soil with a
shallow ground water level shows higher liquefaction susceptibility
than deep ground groundwater level (Kramer, 1996;Kavazanjian
et al., 1997;Bourenane et al., 2018).
CSR, CRR, and FS
The CSR, CRR, and FS values were estimated to evaluate the
liquefaction potential at five selected sites in Hawassa town. The
liquefaction potential analysis for these sites was simulated with
Mw 6.69 of Northridge earthquake ground motion. Therefore,
based on the FS values, each liquefiable layer was delineated. At
site 1, it was evident that the groundwater level (GWL) from the
BH data were observed at a depth of 2 m. The Vs., CSR, CRR, and
FS were determined up to a depth of 30 m. The plots of Vs., CSR,
CRR, and FS versus depth are shown in Figure 10A. It was
observed that, the Vs.for this site ranges from 168 m//s to 542 m/
s. The V
S
was found to decrease up to 5 m before it started
increasing up to a depth of 25 m. In addition, the Vs. showed a
decrease in values at a depth >25 m. The CSR value was varied
from 1.29 to 1.45. Therefore, the CSR values were greater than the
FS limit. The value of CRR ranged from 0.007334 to 1.755. CRR
value was less than FS limit as it goes to a depth of 9.5 m whereas
CRR values were found to be greater than the FS limit when the
depth is greater than 11.9 m. The FS values range from 0.00543 to
1.285. In addition, the FS was <1 when its depth ranges from 2 m
up to 11.9 m depth while the FS was >1 at a depth starts from
11.9 m up to 30 m. Finally, the CSR, CRR, and FS values of the
site have shown that liquefaction is expected from a depth range
of 2 m up to 11.9 m, indicating the site was found to be
liquefiable.
FIGURE 9
Liquefaction susceptibility criteria curves of silty sandy soil in the study area.
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According to the results of site 2 (Figure 10B), the GWL from
the BH data were found at a depth of 22.5 m, the Vs., CSR, CRR,
and FS profile plot with a depth have shown that Vs. values range
from 269 m/s to 452 m/s, CSR from 0.00708 to 0.016, and CRR
from 0.229 to 1.361. The Vs. values decreased at a depth of 10 m
before they started to increase. The FS at this site was greater than
one that means it ranged from 17.38 to 114.29. According to the
values of CSR and CRR, CSR was less than CRR. As a result, the
FS value exceeds one. The values of FS at this site have shown that
there is no liquefaction hazard at the given earthquake motion.
From the BH data at site 3 (Figure 10C), the GWL was
encountered at a depth of 27 m. As it was seen from the
profile plot of depth versus Vs. it varies from 233 m/s to
390 m/s. This profile analysis shows that (i) the Vs. decreases
from the surface to a depth of 9.5 m, increases at a depth of 11 m,
and decreases at a depth of 30 m, (ii) the CSR values varies from
0.07749 to 0.0165, (iii) CRR values varies from 0.406 to 1.052, and
(iv) FS ranges from 52.0407 to 63.41. Based on the comparison of
CSR, CRR, and FS values, CRR higher than CSR and thus FS was
greater than one. When the FS was >1for this site, then
liquefaction will not be induced for the given earthquake
ground motion. From the BH data at site 4, the GWL was
encountered at 27.5 m. The Vs., CSR, CRR, and FS profile
versus depth were presented in Figure 11A. The Vs. profile
has been exhibited from 241 m/s to 284 m/s. Furthermore, the
Vs. were increased at a depth of 9.5 m. After this depth, the Vs.
have shown decreasing or relatively constant. The values of CSR
ranged from 0.00442 to 0.0203, CRR from 0.292 to 0.481 and the
FS from 23.5 up to 66.2. The analysis of CSR and CRR has shown
that CRR values are greater than CSR. Due to the higher value of
CRR, the FS was found to be greater than one, indicating that this
site doesn’t have any liquefaction problem. As it was seen from
the results of site 5 in Figure 11B, the GWL was found at a depth
of 24.8 m. The plot of Vs. value of this site ranges from 291 m/s to
444 m/s. In addition, its profile has shown that Vs. increase from
the ground surface to a depth of 30 m. The CSR varied from
0.051 to 0.0123 while the CRR ranges from 0.423 up to 1.26. The
earthquake resistance force at this site was greater than that of the
driving cyclic stress at this site. Since the FS is >1, the liquefaction
problem doesn’t not exist on this site. Generally, the liquefaction
potential is predicted where there is possibility of increase in pore
water pressure and loss of shear strength occurred in a soil.
Hence, the liquefaction potential was predictable at site 1
(Figure 10A) due to increase in pore water pressure and loss
FIGURE 10
CSR, CRR, FS, and FS limit (A) at site 1, (B) at site 2 and (C) at
site 3.
FIGURE 11
CSR, CRR, FS, and FS limit (A) at site 4 and (B) at site 5.
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of shear strength occurred in a soil (Ganapathy and Rajawat,
2012;Setiawan and Jaksa, 2018;Subedi and Acharya, 2022), but
the soil at site 2 as shown in Figure 10B and site 3 at Figure 10C,
site 4 at Figure 11A and site 5 at Figure 11B was not predictable as
they had no chance of pore water pressure and loss of shear
strength for the given ground motion (Naik et al., 2020). Thus,
the result of this study was supported by the findings of Ji et al.
(2021).
LPI, PL, and PG
The LPI was evaluated for the selected sites to check the site
probability of each liquefied layer. The LPI and PL were
estimated for the sites by FS, whereas PG was determined for
the five sites in Hawassa town based on the LPI values. The LPI
values for the sites in Hawassa town have a range from 0 up to
54.16 as shown in Table 4. For the selected sites, the LPI values
ranged from very 0 to 54.16 liquefaction potential categories.
TheLPIvaluewasfoundtobe54.16forsite1,0forsite2,site
3, site 4 and site 5. In addition, the liquefaction potential
category for site 1 was found to be 54.16 and 0 for the site 2,
site 3, site 4, and site 5. The PL and FS of site 1 varied from
0.21 to 1 and 0.00543 to 1.285, respectively. The PL for site 2,
site 3, site 4, and site 5 were found to be less than 0 while the FS
of site 2, site 3, site 4 and site 5 were greater than 23. Based on
the classification of the PL and FS, site 1 is almost certain that
it will liquefy for Mw 6.69. The PG value for site 1 varied from
0.0073 to 0.993, while the PG for site 2, site 3, site 4, and site
5 was nearly zero.
The description of the risk of PG values for site 1 indicated a
range from low to extremely high, while the description of the
risk of PG values for site 2, site 3, site 4, and site 5 falls from
extremely low to none. The LPI values for site 1 (Figure 12) has
shown an increases from the surface to a depth of 9.3 m,
decreases at a depth of 11.5 m and finally goes to zero from
14.6 to 30 m. According to the LPI values, only site 1 has very
high liquefaction potential category but site 2, site 3, site 4, and
site 5 have no liquefaction potential. The PG at site 1 will not
cause ground failure at a depth greater than 11.5 m (Figure 12).
As a result, the likelihood of liquefaction ground failure at a
shallow depth should be considerable. The PL (Figure 12) will be
higher at a shallow depth up to 10 m and then decrease from
11.6 to 30 m. The results of PL and PG for site 1 showed very high
but for site 2, site 3, site 4, and site 5 very low values similar to
Bahari et al. (2020).
Conclusion
In the paper entitled “Site response and liquefaction hazard
analysis of Hawassa town, Main Ethiopian Rift”, the ELA,
liquefaction susceptibility criteria, FS, LPI, PL, and PG, which
are most widely used in earthquake engineering, were employed.
The main conclusions are summarized as follows:
✓The Vs.30 values ranged from 248.9 m/s to 371.3 m/s. As a
result, the study area falls into C and D seismic site classes
on the accordance of NEHRP (2005) but into B and C based
on the EC 8 (2003) code.
✓The liquefaction susceptibility criteria demonstrated that
the soils in the study area are highly susceptible to
liquefaction since all the soil samples in the study area
fall at the boundaries of most liquefiable soil.
✓The ground motion prediction parameters like PGA (g), 5%
damped SA (g), and maximum strain (%) were estimated
for the study area in order to predict the potential of
liquefaction for the earthquake of Mw6.69 with a PGA of
0.11 g using the DeepsoilV.7 program.
✓The PGA (g) ranged for the sites at the ground surface from
0.166 to 0.281 g. So, the PGA (g) value is amplified at the
shallow depth due to the local site effects.
✓The maximum strain (%) has shown that the ground
deformation reaches its maximum strain (%) at a
shallow depth for the selected sites.
✓The large value of 5% damped SA (g) occurred from 0.1 to
1 s but reached maximum values at 0.4 s. As a result, the
engineering structure damage will be greater at selected
TABLE 4 LPI and liquefaction potential categories for site 1, 2, 3, 4 and
5 at Hawassa town.
Serial number Site ID number LPI Liquefaction
potential category
1 1 54.16 Very high
2 2 0 Very low
3 3 0 Very low
4 4 0 Very low
5 5 0 Very low
FIGURE 12
FS, LPI, PL and PG at site 1.
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sites when the period reaches 0.4s. Therefore, the Hawassa
town needs site-specific seismic design in the period range
between 0.1 and 1 s due to local site effects to minimize the
settlement and deformation of civil engineering structures.
✓FS value ranged from 0.00543 to 114.29. As a result, only
one site is likely to liquefy due to high pore water pressure
and loss of shear strength in the soil, but the other sites are
unlikely to liquefy with similar conditions.
✓LPI demonstrated that only one site has a very high
probability of liquefaction due to high liquefaction
potential values.
✓The PL results also showed that only one site will liquefy at
Mw 6.69, but the other typical sites will not liquefy.
✓Generally, the FS, LPI, PL, and PG values have shown that
there is a site that will be susceptible to liquefaction and
ground settlements at a shallow depth. The site which is
more susceptible to liquefaction hazards is due to increased
pore water pressure and loss of shear strength in the soil.
✓The results from this study suggest that the Ethiopian
seismic code (ES EN 1998:2015) should be revised and
updated by incorporating local site effects into the current
seismic code to minimize earthquake hazard in the region.
Moreover, enforcing the Ethiopian building codes will also
reduce the susceptibility of the town.
✓The results of this study and others revealed that
comprehensive study of ELA and Vs-based liquefaction
hazard analysis helps better understand the site conditions
for seismic microzonation, earthquake mitigation and
prevention purposes, and earthquake resistant design.
✓The concerned government bodies at the town, region, and
federal levels should take tangible action to mitigate the
earthquake induced-liquefaction problems by densification
of likely highly liquefiable sites. In general, the results from
this study will provide important information for
researchers, engineering designers, earthquake engineers,
and planners to minimize earthquake induced-liquefaction
hazards in Hawassa town.
Data Availability Statement
Some of the data analyzed during this study was included in
this article. The remaining datasets used for this study are
available from the corresponding author and can be accessed
upon reasonable request.
Author contributions
For the completion of this manuscript, AA was collected,
processed, compiled, analyzed, and simulated the site response
data using the DeepsoilV.7 program, liquefaction analysis, and
write-up. MM and KW have critically reviewed the draft
manuscript and have done technical editing and English
correction in order to improve the overall quality of the
manuscript. All the authors read and approved the final manuscript.
Acknowledgments
The author would like to thank both supervisors for their
guidance and supervision on this research. He also would like to
extend his gratitude to acknowledge ARCON Design Build plc
for providing geotechnical data. The SDCSE would also be
thanked for providing borehole data. Without the
DeepsoilV.7 software and the Pacific Earthquake Engineering
Research Center (PEER) database (http://peer.berkeley.edu), this
article would not have been realized. Hence, they should be
acknowledged for this important role.
Conflict of interest
The authors declare that the research was conducted in the
absence of any commercial or financial relationships that could
be construed as a potential conflict of interest.
Publisher’s Note
All claims expressed in this article are solely those of the
authors and do not necessarily represent those of their affiliated
organizations, or those of the publisher, the editors, and the
reviewers. Any product that may be evaluated in this article, or
claim that may be made by its manufacturer, is not guaranteed or
endorsed by the publisher.
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