Evaluation of liquefaction potential by standard penetration test and shear wave velocity methods: a case study

Abstract

Dynamic loads such as earthquakes can cause major failure to soil, and the foremost of these failure is soil liquefaction. In engineering studies, liquefaction analyses can be performed by different methods based on the results of field and laboratory experiments. In this study, the liquefaction potential of the soils of the Batman settlement zone, located close to the East Anatolian Fault Zone in Türkiye, was evaluated. In the study area, Meydan site consists of low-medium plasticity clay (CL) at a depth of 1.5–3 m and silty sand (SM) at a depth of 3.5–12.45 m; Bahçelievler site consists of low-medium plasticity clay (CL) at a depth of 1.5–7 m and silty sand (SM) at a depth of 7–15 m. For liquefaction analysis, magnitudes and accelerations of design earthquake were selected as 0.30 g for accelerations, and 7.5 and 6.5 for magnitudes. For this purpose, the effects of soil and earthquake parameters on soil liquefaction were evaluated using the standard penetration test (SPT) and shear wave velocity (Vs) methods with the Microsoft Excel-based SoilEngineering program and the obtained results were correlated and discussed. It is emphasized in this work that liquefaction potential analyses using soil and earthquake parameters provide more reliable results. In addition, the soil strata of locations where the liquefaction risk potential is high were found to have total settlement of approximately 36.87 to 36.2 cm, parallel to the high liquefaction risk, and it was determined that there may be high settlement in the area.

1 Introduction

Earthquakes, which can be defined as dynamic loads that disrupt the equilibrium of the soil, cause cyclic loading. The behavior of soil against these cyclic stresses varies depending on soil type, physical properties, earthquake characteristics, dynamic behavior, and stress-displacement properties. Therefore, cyclic loads cause deformations of different sizes in each soil type (Koc 2007). And then soil failure may lead to damage because of liquefaction. Liquefaction is the process in which soil changes from a solid phase to a liquid phase, often as a result of cyclic loads during earthquakes that decrease effective stress and increase pore water pressure in water-saturated soil strata.

The term “liquefaction” was coined by Mogami and Kubo in 1953 (Kramer 1996), and the devastating effects of liquefaction on soil and local structures during earthquakes, such as the Alaska (M_w = 9.2) and Niigata (M_w = 7.5) earthquakes in 1964, have drawn significant attention to this phenomenon.

As a result, extensive field and laboratory experiments have been conducted to investigate the physics and mechanisms of liquefaction. The first major studies in this field include “Liquefaction of Saturated Sands During Cyclic Loading” by Seed and Lee (1966), “An Experimental Study of Liquefaction of Saturated Sands” by Yoshimi (1967), and “Characteristics of Liquefied Sands During Mino-Owari, Tohnankai, and Fukui Earthquakes” by Kishida (1969).

Several researchers, including Castro (1969), Seed and Idriss (1971), Seed (1976) and Stokoe et al. (1988), have conducted studies based on field measurement data obtained from the standard penetration test (SPT), cone penetration test (CPT), and shear wave velocity (Vs) methods for laboratory studies and liquefaction analysis. After the 1964 Niigata and Alaska earthquakes, Seed and Idriss developed a method for calculating soil liquefaction potential, which they called the “simplified method,” in 1971. Their method was subsequently improved upon by Seed et al. (1975a, b, 1983, 1985) and Youd and Idriss (2001). In the early 2000s, further studies were conducted to refine this approach, including work by Youd et al. (2001), Cetin et al. (2004b), Idriss and Boulanger (2008), and Boulanger and Idriss (2014). These studies have contributed to significant improvements in the calculation of soil liquefaction potential.

Using SPT-N values and Vs data, which are two significant parameters in determining the risk of liquefaction, many researchers have developed various empirical formulas to be applied (Dikmen, 2009; Akın et al. 2011). In addition, many researchers have performed liquefaction analysis using SPT-N values and Vs data (Ulamis and Kilic 2008; Karakas and Coruk 2010; Ulamis and Kilic 2012; Akkaya et al.2018a; Tunusluoğlu and Karaca,2018; Eyisüren et al. 2023). Researchers have performed several studies to determine liquefaction potential using Vs, which is a significant parameter in evaluations of the dynamic properties of soils (Dobry et al. 1981; Tokimatsu and Yoshimi 1983; Andrus and Stokoe 2000;Uyanık 2002; Özçep et al. 2003; Zarif et al. 2004; Zarif et al. 2005; Özçep and Zarif 2006; Uyanık, 2006; Dadashpour et al. 2009; Özçep and Zarif 2009; Uyanık and Taktak 2009; Kayen et al. 2013; Uyanık et al.2013a; Özçep et al.2014; Porcino and Tomasello 2019; Uyanık2020; Duan et al. 2023).

In this study, liquefaction analyses were performed using the Microsoft Excel-based SoilEngineering program (Özçep, 2010) with SPT and Vs field data obtained in Batman Province, which is located close to the East Anatolian Fault Zone in Türkiye. Batman Province, located in the Southeastern Anatolia Region of Türkiye, has various soil problems and is also experiencing rapid construction development. Thus, the effects of earthquake parameters and soil properties are compared and discussed in this study based on liquefaction analyses conducted for Batman Province. In addition, the importance of evaluating soil properties and earthquake parameters together in liquefaction potential analyses is emphasized.

2 Materials and methods

The first simplified studies on evaluating the liquefaction potential of soil through stress were conducted by Seed and Idriss (1971). In their work, the safety of the soil in the face of liquefaction during an earthquake was calculated by comparing the cyclic resistance ratio (CRR) of the soil to the cyclic stress ratio (CSR) resulting from the earthquake. The CSR during an earthquake as described by Seed and Idriss (1971) is shown in Equ. 1.

$$CSR = 0.65*(a_{{\max }}/g) (\frac{{\sigma _{0} }}{{\sigma^{{\prime}}_{0} }}) *r_{d}$$

(1)

In Equ. 1, g represents gravitational acceleration; σ₀ and σ₀ signify the overburden pressure and effective overburden pressure at the depth (z) at which the liquefaction analysis was conducted, respectively; and r_d is the stress reduction factor, which is determined by the depth-based expressions given in Equ. 2 and 3 (Liao and Whitman 1986).

$$r_{d} = 1.0 - 0.00765*{\text{z}}\quad {\text{z}} \le 9.15{\text{m}}$$

(2)

$$r_{d} = 1.174 - 0.0267*z\quad 9.15m{\text{ }}z \le 23\;{\text{m}}$$

(3)

2.1 Liquefaction analysis according to the Turkish Building Earthquake Code (TBEC 2018)

The Turkish Building Earthquake Code (TBEC, 2018) aimed to gather earthquake data for four different levels of earthquake ground motion and earthquake parameters based on the “Earthquake Hazard Map of Türkiye” (EHMT) (https://tdth.afad.gov.tr/). In accordance with the TBEC, the peak ground acceleration (PGA), which represents the largest horizontal ground acceleration value, was determined for the site based on the seismic hazard map values at the location coordinates. This map can be accessed through the website of the Disaster and Emergency Management Presidency (AFAD). To this end, four earthquake hazard maps were prepared for four different levels of earthquake ground motion, including two dimensionless spectral acceleration maps and two maps containing spectral acceleration coefficients for short period map spectral acceleration (S_S) and spectral acceleration coefficients for a period of 1.0 s (S₁). The map spectral acceleration coefficients were converted to design spectral response acceleration (S_DS and S_D1) by employing local effect coefficients chosen based on the local soil class, and horizontal elastic design spectra were derived from these coefficients (Fig. 1).

Fig. 1

Horizontal elastic design spectrum (TBEC 2018)

The determined map spectral acceleration coefficients S_S and S₁ were converted to design spectral response accelerations ${\mathrm{S}}_{\mathrm{DS}}$ and ${\mathrm{S}}_{\mathrm{D}1}$ as shown in Equ. 4 and Equ. 5 (TBEC, 2018).

$${\text{S}}_{{{\text{DS}}}} {\text{ = S}}_{S} \; \cdot \;{\text{F}}_{s} {\text{ }}$$

(4)

$${\text{S}}_{{{\text{D1}}}} {\text{ = S}}_{1} \; \cdot \;{\text{F}}_{1}$$

(5)

In Equ. 4, ${S}_{S}$ represents the short period spectral acceleration value and ${F}_{S}$ is the local ground effect coefficient for the short period region. ${\mathrm{S}}_{\mathrm{DS}}$ is the short period (0.2 s) dimensionless design spectral response acceleration. In ${\mathrm{S}}_{\mathrm{DS}}$, “S” stands for spectral acceleration, the subscript “D” stands for design, and the subscript “S” stands for short. Local ground effect coefficients F_S and F₁ are given in Tables 1 and 2.

Table 1 Local soil impact coefficient (Fs) for the short period zone (TBEC 2018)

Table 2 Local soil impact coefficient (F₁) for the short period zone (TBEC 2018)

Horizontal design spectrum corner periods were defined based on ${T}_{A}$ and ${T}_{B}$ with ${S}_{DS}$ and ${S}_{D1}$, as given in Equ. 6 and Equ. 7. The transition period to the constant displacement region was accepted as ${T}_{L}$ = 6 s (TBEC, 2018).

$$T_{A} = 0.2\frac{{S_{D1} }}{{S_{DS} }}$$

(6)

$$T_{B} = \frac{{S_{D1} }}{{S_{DS} }}$$

(7)

Within the framework of the Turkish Building Earthquake Code (TBEC, 2018), it was determined during SPT analyses performed in the boring stage that some physical factors had effects on the test and corrections for those factors were proposed to ensure that the test results could be evaluated in line with a common standard. The blow counts measured for a field during SPT-N analysis depend on many factors including overburden pressure, borehole diameter, rod length, energy ratio, and type of sampler. For this reason, the SPT-N values to be used in calculations based on SPT should be corrected first (Table 3). In the present study, the corrected SPT-N values for an overburden pressure of 100 kPa and energy efficiency of 60%, N_1,60, were obtained from Equ. 8.

$$N_{1,60} = N_{60}*C_{N} \Rightarrow N_{60}=N_*C_{R}*C_{S}*C_{B}*C_{E}$$

(8)

Table 3 SPT correction factors (TBCE 2018)

The left-hand side of Equ. 8 represents the corrected SPT blow count, $N$ is the measured number of SPT blow counts in a given field, ${C}_{N}$ is the depth correction factor, ${C}_{E}$ corresponds to the hammer energy ratio (ER) correction factor, ${C}_{B}$ is the borehole diameter correction factor, ${C}_{R}$ denotes the rod length correction factor, and ${C}_{S}$ is a correction factor for samplers with or without liners.

The curve separating the regions where liquefaction occurs and where it is not observed in cases of earthquakes of magnitude Mw = 7.5 for clean sands with less than 5% fine grains (FC ≤ 5%) was defined as the clean sand curve by Youd et al. (2001) (Fig. 2). The cyclic resistance ratio (${CRR}_{7.5}$) for this condition can be formulated as shown in Equ. 9.

$${\text{CRR}}_{7.5}=\frac{1}{34-N_{1,60f}}+\frac{N_{1,60f}}{135}+\frac{50}{\left[10N_{1,60f}+45\right]^{2}}-\frac{1}{200}$$

(9)

Fig. 2

Corrected SPT (N)-CSR chart in clean and silty sand for an earthquake with a magnitude of 7.5 (Youd et al. 2001)

The Turkish Building Earthquake Code (TBEC, 2018) also addresses the determination of liquefaction potential according to shear wave velocity (Vs), which is one of the dynamic properties of soils. Vs values can be measured using classical geophysical methods such as seismic refraction and seismic reflection, as well as borehole geophysical methods such as cross-hole, down-hole, and seismic CPT or analysis methods such as multichannel analysis of surface waves (MASW) and spectral analysis of surface waves (SASW).

Liquefaction resistance depending on Vs as obtained by seismic experiments was calculated based on the formula of Andrus and Stokoe (2000) as shown in Equ. 10.

$${\text{CRR}}_{7.5} = 0.022\left[\frac{(V_{s1})}{100}\right]^{2}+{2.8}\left[\frac{1}{215-(V_{s1})}-\frac{1}{215}\right]$$

(10)

In this method, the S-wave velocity (${V}_{S1}$) must first be corrected in order to be used in liquefaction calculations (Fig. 3). Here, corrected S-wave velocity (${V}_{S1}$) was calculated using the overburden load and Vs relative to vertical stress as shown in Equ. 11.

$$V_{S1} =V_{S} \left( {P_{a} /\sigma_{0}^{\prime } } \right)^{0.25}$$

(11)

Fig. 3

Corrected Vs-CSR Chart for clean and silty sands for an earthquake with a magnitude of 7.5 (Andrus and  Stokoe 2000)

In the equation above, ${P}_{a}$ is atmospheric pressure of 100 kPa and σ′v is the effective vertical stress measured in kPa.${V}_{S1}^{*}$ is the upper limit of shear wave velocity at which liquefaction occurs depending on the fine grain ratio (FC). In the equation for sandy soil, ${V}_{S1}$ is as follows:

$${\text{CRR}}_{7.5} = 0.022\left[\frac{(V_{s1})}{100}\right]^{2}+{2.8}\left[\frac{1}{215-0.5(FC-5)-(V_{s1})}-\frac{1}{215-0.5(FC-5)}\right]\quad \quad 5\% < F.C. < 35\%$$

(12)

$${\text{CRR}}_{7.5} = 0.022\left[\frac{(V_{s1})}{100}\right]^{2}+{2.8}\left[\frac{1}{200-(V_{s1})}-\frac{1}{200}\right]\quad \quad F.C. \gg 35\%$$

(13)

2.2 Liquefaction safety conditions

The liquefaction potential of a soil profile is assessed by comparing its loading and resistance. The safety factor is determined by considering the cyclic resistance ratio (CRR) of the soil against liquefaction and the cyclic stress ratio (CSR) induced by the earthquake. Equation 2.19 presents the safety factor calculated using the corrected SPT-(N1)60 values proposed by Youd et al. (2001) and Idriss and Boulanger (2008). Moreover, Equ. 14 incorporates the effect of the magnitude scaling factor (MSF) on the safety factor against liquefaction.

$$FS = \frac{CRR}{{CSR}}\text{MSF}$$

(14)

As a general rule, no liquefaction is expected for safety factors greater than 1.2, potential liquefaction is expected for safety factors between 1 and 1.2, and liquefaction is expected for safety factors less than 1 (Seed and Idriss, 1971). To consider the impact of duration effects (number of loading cycles) on the initiation of liquefaction, the magnitude scaling factor (MSF) is employed. Figure 4 illustrates the relationships between the MSF factors (Idriss and Boulanger 2010), which are determined using Equ. 15. Idriss (1999) developed this earthquake magnitude correction factor.

$${\text{MSF}} = 6.9 {\text{exp}}\left( {\frac{{ - M_{w} }}{4}} \right) - 0.0{6}\quad \quad M_{w} > 5.2$$

(15)

Fig. 4

Magnitude scaling factor (MSF) relationships (Idriss and Boulanger 2010)

According to the Turkish Building Earthquake Code (2018), the safety factor against liquefaction is to be calculated with Equ. 16. If the condition given in Equ. 16 is not met, then liquefaction is to be expected, and if the condition is met, then liquefaction is not expected.

$$FS_{L} = \frac{{\tau_{R} }}{{\tau_{earthquake} }} \gg 1.10$$

(16)

In Equ. 16, ${\tau }_{R}$ and ${\tau }_{earthquake}$ refer to liquefaction resistance and the average cyclic shear stress caused by the earthquake on the soil, respectively.

2.3 Soil settlement due to liquefaction

The generation of excessive pore water pressure during an earthquake can result in cyclic soil movements during and after the event. The dissipation of pore water pressure following the earthquake leads to soil volume changes. Such volumetric strains, which cause soil settlement due to liquefaction, can impact the performance of structures and cause strains on their bases. Empirical studies by Lee and Albaisa (1974) and Yoshmini (1975) investigated the volumetric changes caused by excessive pore water pressure dispersion in saturated sands under cyclic loads in laboratory conditions. More recent related studies proposed empirical maps for estimating volumetric strain in liquefied clean sands. These include Tokimatsu and Seed (1987), Ishihara and Yoshimine (1992), Shamoto (1998), Pradel (1998), Unutmaz and Çetin (2005), and Bilge and Çetin (2006). These maps were developed based on dynamic triaxial compression tests on sands of varying relative densities. Ishihara and Yoshimine (1992) developed a method that evaluates the volumetric strain of each layer separately in calculations for soil settlement after an earthquake. The calculation for soil settlement after an earthquake is given in Equ. 17. The volumetric strain can be determined using the graph in Fig. 5, which depends on the safety factor (SF) and the relative density in soil layers (Dr).

$$S_{sat} = (\varepsilon_{V} /100)*z$$

(17)

Fig. 5

The relationships between safety factor, F_S, and post-liquefaction volumetric strain ${\varepsilon }_{V}$ (Ishihara and  Yoshimine 1992)

In Equ. 17, z is the thickness of the layer in meters, ${\varepsilon }_{V}$ is the volumetric strain, and S_sat is the corresponding saturated soil layer.

2.4 SoilEngineering: a Microsoft Excel® spreadsheet program

In recent years, many programs that provide computer-aided solutions to analyze liquefaction potential through the use of different parameters have been developed in the field of geotechnical engineering. Some of these programs include LIQUITER, which works based on the method of Seed and Idriss (1971) and enables calculations of safety factors against liquefaction for saturated cohesionless soils subjected to earthquake loads; LiqIT, which enables liquefaction analysis of saturated loose soils; SPTLIQ, which enables calculations of liquefaction potential and settlement caused by earthquakes; LiquefyPro, which enables evaluations of liquefaction potential and earthquakes; and CUMLIQ, which enables calculations of the liquefaction potential of alluvial soils (Tolon, 2013). In addition, soiLique, developed by Bekin and Özçep (2021), offers the first deterministic soil liquefaction analysis method based on MATLAB that includes a graphical user interface (GUI). The fundamental characteristic of soiLique is its independence from MATLAB®. In other words, soiLique can be used without MATLAB® by installing MATLAB® Runtime Environment v9.7 to run. In the present study, liquefaction analysis calculations were performed using the SoilEngineering program developed by Özçep (2010). It is an open coded software. SoilEngineering is a free Microsoft Excel-based program that is widely used for the geotechnical and geophysical analysis of soils, especially in Türkiye. Its application in geotechnical and geophysical analysis of soils under static and dynamic loads is very practical. Analysis of soil problems in geotechnical research can be easily performed using this program. In addition, this program also has the advantage of calculating data numerically and presenting the results with visual graphs (Figs. 6, 7).

Fig. 6

SoilEngineering Microsoft Excels Program overview

Fig. 7

SoilEngineering Microsoft Excels Program Main Options view

3 Geological setting

Batman Province is located in the Southeastern Anatolia Region of Türkiye. The approximate area of the province is 4,477 km² and it is located at an altitude of 550 m. The oldest stratigraphic unit observed in and around Batman Province is the Hoya Formation of the Midyat Group, which consists of a Lower Eocene-Lower Oligocene sparse argillaceous limestone level with limestone, dolomitic limestone, and dolomites. This formation is concordant with the Gercüş Formation near Hasankeyf, located in the vicinity of Batman Province (Tuna 1973; Açıkbaş and Baştuğ 1975; Perinçek 1979). The Hoya Formation is composed of massive and medium to thickly layered, sometimes thinly to medium layered, cream, beige, gray, light gray, black, or white nummulite, miliolid, algal, echinoid, gastropod, lamellar, shell-sediment, hard limestones with sporadic dolomitic limestones. This stratigraphic unit, consisting of white, red, coffee-colored, green, beige, and gray gypsum, anhydride, shale, dolomite, and sandy alluvion dolomites, was first identified by Bolgi (1961). The Germik Formation, consisting of Oligocene-aged argillaceous limestones and sporadic dolomitic limestone, overlies the Hoya Formation. While the Germik Formation concordantly overlies the Hoya Formation, it is nonconcordantly overlain by the Şelmo Formation. The Germik Formation consists of white, beige, gray, and dirty yellow soft gypsum, anhydride, shale, and dolomites that are not resistant to dispersive physical conditions. The Germik Formation’s thickness is no more than 465 m (Eren et al. 2012). These formations are nonconcordantly overlain by the Şelmo Formation, which is composed of Middle-Upper Miocene conglomerate, sandstone, and mudstone. The Şelmo Formation constitutes a large part of the study area located in the city center of Batman (Eren et al. 2012). The Quaternary-aged basalts forming Kıra Mountain overlie the Şelmo Formation, and Quaternary units consisting of alluviums surface around the Batman Creek Valley and Batman (Fig. 8).

Fig. 8

Geologic map of Batman City and surrounding area, modified from The Institute of Mineral Research and Exploration (MTA) 1:500.000 scale geological map (Sunkar and Tonbul 2013)

3.1 Seismotectonics and seismic hazard of the study area

The study area and its immediate surroundings are located to the south of the “trigger zone” or contact zone of many earthquakes that occur in Türkiye. In the study area, located to the south of the edge fold region, there are parallel and regular fold areas that occur as a result of north–south directional compression movement. In order to understand the geological features of the study area, it is necessary to understand the relationship between the Anatolian Plate and the Arabian Plate, which greatly affects the geological structure of the region, and to discuss the tectonic features of a more extensive area. The Southeastern Anatolia Region and its surroundings, where the study area is located, are sheared by the East Anatolian Fault Zone (EAFZ), the Bitlis-Zagros Suture Zone (BZSZ), and the Dead Sea Fault Zone, which are among the largest tectonic structures of Türkiye, and by smaller conjugated faults of those major faults (İmamoğlu and Çetin, 2007). The EAFZ, which is one of the two main fault belts that are most active in Türkiye, spans over 580 km between Karlıova and Antakya and plays an important role in the geodynamic evolution and seismicity of the region (Arpat and Şaroğlu 1972, 1975; McKenzie 1972; Seymen and Aydın 1972; Ambrasseys, 1989; Taymaz et al. 1991). The Arabian Plate collides with the Anatolian Plate along the BZSZ, which is a region of deformation, causing the Anatolian Plate to move westward (Eyidoğan 1983; Barka and Reilinger 1997). As a result of the African-Arabian and Eurasian plates approaching each other on the north–south line, the ancient Tethys Sea, which is considered to have been the ancient ancestor of the Mediterranean, was closed off in the Lower-Middle Maastrichtian period. Following this, the Eurasian and Arabian plates collided along the Southeast Anatolian Thrust or BZSZ in the Tortonian period, about 10 million years ago (Şengör, 1980). The Anatolian Plate moved west under the influence of the left-lateral strike-slip EAFZ and the right-lateral North Anatolian Fault Zone (NAFZ) (Barka and Kadinsky-Cade 1988; Şaroglu et al. 1992) (Fig. 9). The East Anatolian Fault (EAF) starts from the Karlıova junction point in the northeast and continues toward the Türkoğlu junction in the southwest. The EAFZ starts from the Karlıova triple junction with the NAFZ in the north and connects to the Dead Sea Fault System in the SW. The length of the EAFZ, the general direction of which is NE-SW, is about 580 km. The EAFZ is a left-lateral strike-slip fault and consists of 6 segments (Fig. 10) (Şaroğlu et al. 1992a)

Fig. 9

NAFZ, EAFZ, Massives, Terranes, BZKK, and other important tectonic structures in the Geology of Turkey (adopted from Marroni et al. 2020)

Fig. 10

Eastern Anatolian Fault Zone segments (adopted from Demirtaş and Erkmen 2019)

Deterministic and probabilistic approaches are used to conduct seismic hazard analyses, especially in the case of developing ground movements. Seismic hazard analysis is defined as the determination of ground movements caused by earthquakes in a certain location and time period (Erdik et al. 1999). Deterministic and probabilistic approaches are widely used for the definition and evaluation of ground motion. Earthquake hazards calculated deterministically represent the level of ground movement that the largest earthquake that may occur in a region will create regardless of the time dimension (Seismic Microzonation for Municipalities: Manual, 2004). The active fault zones near the study area and relevant earthquakes in historical and instrumental periods are shown in Fig. 11. For deterministic seismic hazard analysis, two fault models within the EAFZ, namely the A fault (fault rupture is 16 km) and B fault (fault rupture is 120 km), were selected (Table 4 and Table 5). The levels of ground motion for magnitudes of 7.5 and 6.5 are shown in Table 6. Table 6 also provides estimated accelerations for the region, calculated using several different attenuation relationships. It is analyzed that the results calculated for the region support each other.

Fig. 11

Active fault zones, earthquakes in historical and instrumental periods near study area (map is redrawn by Erdik et al 1999)

Table 4 Selected fault model (A: fault rupture length is 16 km) within East Anatolian Fault Zone

Table 5 Selected fault model (A: fault rupture length is 120 km) within East Anatolian Fault Zone.

Table 6 Ground motion level for magnitüde 7.5 and 6.5

3.2 Geotechnical investigation

For soil liquefaction analyses in the city center of Batman, where the study area is located, 4 boreholes were drilled in 2 locations, namely the Meydan and Bahçelievler sites, and MASW measurement studies were carried out for 4 profiles (Fig. 12). It is thought that these areas selected in the city center of Batman have high liquefaction potentials. Figure 13 shows a general view of the study area. The boreholes drilled in the Meydan site (BH-1 and BH-2) are in an area composed of low-medium plasticity clay (CL) at a depth of 1.5–3 m and one of silty sand (SM) at a depth of 3.5–12.45 m, respectively. The boreholes drilled in the Bahçelievler site (BH-3 and BH-4), on the other hand, are in an area composed of low-medium plasticity clay (CL) at a depth of 1.5–7 m and one of silty sand (SM) at a depth of 7–15 m, respectively. As can be seen in Fig. 14 a, b, according to the results obtained from these boreholes, the soils of the region can be classified mainly as low-medium plasticity clay (CL) and silty sand (SM). The SPT values obtained from the boreholes in the study area are 11–18 for the Meydan site at a depth of approximately 3.0–12.00 m and 6–18 for the Bahçelievler site (Fig. 15) at a depth of approximately 7.0–15.00 m. It was further determined that, at a depth of 3–12 m, the average gravel content is 1.3%, sand content is 68.1%, the fine grain ratio (silt + clay) is 30.6%, and PI = 14.4% at the Meydan site. At a depth of 7–15 m, the average gravel content is 1.4%, sand content is 70.2%, the fine grain ratio (silt + clay) is 28.4%, and PI = 15.3% at the Bahçelievler site. According to seismic refraction and MASW test results, Vs values fall within the range of 175–201 m/s at a depth of 3–12 m at the Meydan site and a range of 175–242 m/s at a depth of 3–12 m at the Bahçelievler site (Fig. 15).

Fig. 12

Location of study area a Meydan and b Bahçelievler site drillings and seismic profile

Fig. 13

General views from the study site

Fig. 14

a Meydan and b Bahçelievler SPT results of the study site

Fig. 15

Example SPT and Vs depth change for Bahcelievler site

4 Results and discussions

In the present study, in order to observe the effects of soil and earthquake parameters on soil liquefaction, the SPT and Vs methods and the Microsoft Excel-based SoilEngineering program (Özçep, 2010) were used to make the application of liquefaction analyses more practical. Subsequently, the obtained results were discussed and interpreted. Liquefaction analyses based on the SPT and Vs methods were conducted using different earthquake magnitudes, design earthquake accelerations, groundwater levels (3 and 7 m), and depths (z = 1.8, 3.3, 4.8, 6.3, 7.8, 9.3, 10.8, 12.3, 13.8, 15.3, 16.8, 18.3, 19.8, and 21.3 m). The data input into SoilEngineering (Özçep, 2010) for liquefaction risk analysis of the study site, namely earthquake magnitude, design earthquake acceleration, groundwater level, and short period (0.2 s) design spectral response acceleration (SDS) values, are given in Table 7. In addition, the input data used for the calculation of shear stress $({\tau }_{earthquake}$) as part of the soil liquefaction analysis performed using SoilEngineering (Özçep, 2010) are given in Table 8, Table 9, Table 10, and Table 11. The values shown in bold color are program inputs. The values in bold color throughout the liquefaction analyses performed according to the SPT method are program inputs used for calculations of soil shear stress $({\tau }_{r}$). Safety factors (SF) were obtained by proportioning soil shear stress $({\tau }_{r}$) and earthquake shear stress $({\tau }_{earthquake}$). In addition, in the analyses performed according to the Vs method, the program output states that liquefaction is not expected when the Vs1 value is theoretically greater than Vs1c (maximum: 220 m/s). The input data and SF values of calculations made according to the SPT method ​​are presented in Table 12, Table 13, Table 14, and Table 15, respectively. The input data and SF values of calculations made according to the Vs method ​​are presented in Table 16, Table 17, Table 18, and Table 19, respectively. In the liquefaction analysis of an earthquake of magnitude 7.5 with design earthquake acceleration of 0.30 g and a short period (0.2 s) design spectral response acceleration (SDS) of 0.481 at the BH-1 borehole in the Meydan site, where the groundwater level is 3 m, the SF values calculated for a depth of 3.3–12.3 m according to the SPT and Vs methods (Table 12 and Table 16) are 0.46–0.57 and 0.4202–0.7941, respectively. The fact that the SF values for the Meydan site are less than 1 indicates that there may be a risk of liquefaction in this area. It is seen that SF values are less than 1 in the results of both analyses. This indicates that the results support each other and the risk of liquefaction is high in this area (Fig. 16 a). In the liquefaction analysis of an earthquake of magnitude 6.5 with a design earthquake acceleration of 0.30 g and a short period (0.2 s) design spectral response acceleration (SDS) of 0.481 at the BH-2 borehole drilled in the Meydan site (Table 9), where the groundwater level is 3 m, the SF values calculated for a depth of 3.3–12.3 m according to the SPT-N method (Table 13) are 0.39–0.41 and the SF values calculated for a depth of 4.8–12.3 m according to the Vs method (Table 17) are 0.5647–0.7931. Both analyses showed that there may be a risk of liquefaction since the SF values are less than 1 at the specified depths (Fig. 16 b). In addition, analyses conducted according to the Turkish Building Earthquake Code of 2018 (Table 20) yielded results compatible with the SPT and Vs results at a depth of 4.0–12.3 m, further supporting the possibility of liquefaction.

Table 7 Earthquake and soil data set of the study area

Table 8 Input data table of SoilEngineering liquefaction analysis program for Meydan site BH-1: Cyclic shear resistance (CSR) (D) and shear stress (${\tau }_{earthquake})$ results (M_W = 7.5)

Table 9 Input data table of SoilEngineering liquefaction analysis program for Meydan site BH-2: Cyclic shear resistance (CSR) (D) and shear stress (${\tau }_{earthquake})$ results (M_W = 6.5)

Table 10 Input data table of SoilEngineering liquefaction analysis program for Bahçelievler site BH-3: Cyclic shear resistance (CSR) (D) and shear stress (${\tau }_{earthquake})$ results (M_W = 7.5)

Table 11 Input data table of SoilEngineering liquefaction analysis program for Bahçelievler site BH-4: Cyclic shear resistance (CSR) (D) and shear stress (${\tau }_{earthquake})$ results (M_W = 6.5)

Table 12 Input data table of SoilEngineering liquefaction analysis program showing standard penetration test (SPT) analysis for Meydan site BH-1: Shear resistance ratio (CRR) (Z) and safety factor (SF) results (M_W = 7.5)

Table 13 Input data table of SoilEngineering liquefaction analysis program showing standard penetration test (SPT) analysis for Meydan site BH-2: Shear resistance ratio (CRR) (Z) and safety factor (SF) results (M_W = 6.5)

Table 14 Input data table of SoilEngineering liquefaction analysis program showing standard penetration test (SPT) analysis for Bahçelievler site BH-3: Shear resistance ratio (CRR) (Z) and safety factor (SF) results (M_W = 7.5)

Table 15 Input data table of SoilEngineering liquefaction analysis program showing standard penetration test (SPT) analysis for Bahçelievler site BH-4: Shear resistance ratio (CRR) (Z) and safety factor (SF) results (M_W = 6.5)

Table 16 Input data table of SoilEngineering liquefaction analysis program showing shear wave velocity (Vs) analysis for Meydan site BH-1: Safety factor (SF) results (M_W = 7.5)

Table 17 Input data table of SoilEngineering liquefaction analysis program showing shear wave velocity (Vs) analysis for Meydan site BH-2: Safety factor (SF) results (M_W = 6.5)

Table 18 Input data table of SoilEngineering liquefaction analysis program showing shear wave velocity (Vs) analysis for Bahçelievler site BH-3: Safety factor (SF) results (M_W = 7.5)

Table 19 Input data table of SoilEngineering liquefaction analysis program showing shear wave velocity (Vs) analysis for Bahçelievler site BH-4: Safety factor (SF) results (M_W = 6.5)

Fig. 16

a Meydan site (BH-1) b Meydan site (BH-2) − 4 c Bahçelievler site (BH-3) d Bahçelievler site (BH-4). Safety factor (SF) and depth change graphs according to standard penetration test (SPT) and shear wave velocity (Vs) data

Table 20 According to Turkish Building Earthquake Regulations (2018), liquefaction risk analysis of Meydan site (BH-1 and BH-2) Safety Factor (SF) values

In the liquefaction analysis of an earthquake of magnitude 7.5 with design earthquake acceleration of 0.30 g and a short period (0.2 s) design spectral response acceleration (SDS) of 0.697 at the BH-3 borehole drilled in the Bahçelievler site, where the groundwater level is 7 m, SF values calculated according to the SPT-N method are 0.33–0.75. Since the SF values calculated for liquefaction analyses according to the SPT-N method (Table 14) at a depth of 7.8–15.3 m and according to the Vs method (Table 18) at a depth of 7.8–10.8 m are less than 1, it can be concluded that there is a possible risk of liquefaction (Fig. 16 c). In the liquefaction analysis of an earthquake of magnitude 6.5 with design earthquake acceleration of 0.30 g and a short period (0.2 s) design spectral response acceleration (SDS) of 0.697 at the BH-4 borehole drilled in the Bahçelievler site, where the groundwater level is 7 m, SF values calculated for a depth of 7.8–15.8 m according to the SPT-N method (Table 15) are 0.33–0.62 while the SF values calculated for a depth of 7.8–13.8 m according to the Vs method (Table 19) are 0.4714–0.8173. Thus, the results of liquefaction analysis performed according to the Vs method suggest a risk of liquefaction at depths between roughly 7 m and 13 m, with no risk of liquefaction at deeper depths (Fig. 16 d). In addition, analyses performed according to the Turkish Building Earthquake Code of 2018 (Table 21) yielded results compatible with the SPT and Vs results at a depth of 7.0–14.0 m, supporting the possibility of liquefaction.

Table 21 According to Turkish Building Earthquake Regulations (2018), liquefaction risk analysis of Bahçelievler site (BH-3 and BH-4) Safety Factor (SF) values

Recompression settlement following liquefaction in sand units can occur very quickly with the sudden damping of the pore water pressure due to dynamic loads. Sands tend to be compacted underground due to earthquake tremors and their underground compaction tendency manifests itself on the surface in the form of soil settlement. In this study, surface settlement amounts resulting from vertical strain due to excess pore water pressure being damped were calculated by multiplying and integrating the volumetric strain value [εv (%)] obtained for each depth at the study site with the depth for which they were obtained and then superposing the results (Ishihara and Yoshimine 1992). As can be seen in Table 22, in the analyses conducted for the BH-1 and BH-2 boreholes drilled in the Meydan site, total settlement amounts of 32.06 cm and 49.8 cm were calculated respectively for depths of 3.3–12.3 m, and in the analyses conducted for the BH-3 and BH-4 boreholes drilled in the Bahçelievler site, total settlement amounts of 36.87 cm and 36.2 were respectively calculated for depths of 7.8–15.3 m. Liquefaction potential is high for the boreholes drilled in the Meydan site (BH-1 and BH-2) at depths of roughly 3.3–12.3 m, and accordingly, there may be high amounts of settlement in the soil layers at these depths (Fig. 17a, b). Liquefaction potential observed for the boreholes drilled in the Bahçelievler site (BH-3 and BH-4) at depths of roughly 7.8–15.3 m was also supported by the high amount of soil settlement at those depths (Fig. 17c, d).

Table 22 Changes in the depth and settlement for study area

Fig. 17

Graphs of changes in depth and settlement amounts in a Meydan site (BH-1) b Meydan site (BH-2) c Bahçelievler site (BH-3) d Bahçelievler site (BH-4)

5 Conclusions

In this study, in order to determine soil liquefaction potential in the residential area of Batman Province, located in the Southeastern Anatolia Region of Türkiye, 4 boreholes were drilled in 2 locations, namely the Meydan and Bahçelievler sites, and MASW measurement studies were carried out for 4 profiles. To determine the risk of soil liquefaction, which is one of the most important sources of damage caused by earthquakes to structures in residential areas, a series of analyses based on SPT and Vs data were performed. Within the scope of the research, SPT and Vs parameters obtained from field studies were analyzed together with earthquake parameters in order to obtain SF values reflecting the risk of liquefaction. These SF values were then interpreted and presented. In addition, all obtained results were compared and evaluated in accordance with the SF results calculated for liquefaction risk according to the Turkish Building Earthquake Code of 2018. Analyses of liquefaction risk and liquefaction-related settlement, which are crucial soil problems in geotechnical engineering, were conducted using the computer-aided SoilEngineering program (Özçep, 2010), which provides collective calculation methods and visual graphics. It was found that all analyses performed to determine liquefaction potentials supported each other. Thus, calculating the liquefaction risk according to both soil and earthquake characteristics in future studies will provide more useful and reliable results in solving geotechnical problems.

Liquefaction analyses were carried out in this study using design earthquake features with magnitudes of 7.5 and 6.5, reflecting the earthquakes that may be produced within the EAFZ located close to the study area. In the liquefaction analysis of an earthquake of magnitude 7.5 with design earthquake acceleration of 0.30 g and a short period (0.2 s) design spectral response acceleration (SDS) of 0.481 for the BH-1 borehole drilled in the Meydan site, where the groundwater level is 3 m, the SF values calculated for a depth of 3.3–12.3 m according to the SPT and Vs methods were smaller than 1. This indicates a potential for liquefaction in the area. In the liquefaction analysis performed for an earthquake with magnitude of 6.5, design earthquake acceleration of 0.30 g, and a short period (0.2 s) design spectral response acceleration (SDS) of 0.481 for the BH-2 borehole in the Meydan site, where the groundwater level is 3 m, the SF values calculated for a depth of 3.3–12.3 m according to the SPT-N method were 0.39–0.41, while the SF values calculated for a depth of 4.8–12.3 m according to the Vs method (Table 17) were 0.5647–0.7931. Both sets of analyses showed that there may be a risk of liquefaction since the SF values were smaller than 1 at the specified depths. In addition, analyses performed according to the Turkish Building Earthquake Code of 2018 yielded results compatible with the SPT and Vs data at a depth of 4.0–12.3 m, supporting the possibility of liquefaction.

In the liquefaction analysis performed for an earthquake of magnitude 7.5 with design earthquake acceleration of 0.30 g and a short period (0.2 s) design spectral response acceleration (SDS) of 0.697 for the BH-3 borehole drilled in the Bahçelievler site, where the groundwater level is 7 m, the SF values calculated for liquefaction analyses using the SPT-N method at a depth of 7.8–15.3 m and the Vs method at a depth of 7.8–10.8 m were smaller than 1. Thus, it can be concluded that there is a risk of liquefaction. In the liquefaction analysis for an earthquake of magnitude 6.5 with design earthquake acceleration of 0.30 g and a short period (0.2 s) design spectral response acceleration (SDS) of 0.697 for the BH-4 borehole drilled in the Bahçelievler site, where the groundwater level is 7 m, liquefaction risks were observed at a depth of 7.8–15.8 m according to the SPT-N method and a depth of roughly 7–13 m according to the Vs method, while no risk of liquefaction was seen at deeper depths. In addition, analyses performed according to the Turkish Building Earthquake Code of 2018 yielded results compatible with the SPT and Vs results at a depth of 7.0–14.0 m, supporting the possibility of liquefaction.

Furthermore, the liquefaction risk indices of the studied zone obtained by the SPT method and those obtained by the Vs method were found to be compatible with each other. Liquefaction analyses were conducted with SPT data for areas of Batman Province with liquefaction potential under the conditions of earthquake moment magnitude of Mw = 7.5 and Mw = 6.5 and design earthquake acceleration of 0.30 g, which could be produced by future earthquakes occurring due to the rupture of the EAFZ. Finally, it was determined that there may be large amounts of settlement in soil layers where the liquefaction risk potential is high.