Building A is believed to have been built between 1965 and 1970 adjacent to the shoreline (Onagawa, 2013). It was used as a repair shop of fishing boats in the past but was being used as a commercial store before the tsunami (Onagawa, 2013). The building was submerged by 0.4 m of seawater at high tide for many months after the tsunami due to the residual subsidence from the earthquake. The building was a three-story RC structure with a mat foundation on hard ground, as shown in Figure 2A. Small openings were observed on the face of this building subject to tsunami flow. As shown in Figure 1, this building overturned seaward, but it is expected that the initial failure was landward (consistent with the other buildings) and this building was then moved during tsunami return flow to its final position (Fraser et al., 2013). On the other hand, the building may have been overturned seaward by the receding wave (Onagawa, 2013). Due to the mat foundation, only building self-weight could provide a resisting moment against the overturning moment from hydrodynamic and buoyancy forces.

Building B is believed to have been built between 1955 and 1975 based on its type of foundation (Nikkei BP Company, 2011). It was an accommodation building with a four-story RC structure and a pile foundation, as shown in Figure 2B. The pile foundation had 32 hollow concrete pipe piles with a pile diameter of 20 cm. This building was moved 70 m from its original position. As shown in Figure 2B, some piles were pulled from the ground, and some were broken under the foundation. Twelve of the 32 piles under the foundation appear to have been effective in resisting the overturning moment (Kabeyasawa et al., 2012). No spiral reinforcing bar was observed inside the piles, but six longitudinal reinforcing bars were observed inside each pile. The settled ground around a neighboring building indicated the occurrence of soil liquefaction (Tokimatsu et al., 2012).

Building C is believed to have been built between 1980 and 1985 (Onagawa, 2013). The building was used as an accommodation facility in the past but was being used as offices for private business and accommodations for sailors before the tsunami (Onagawa, 2013). The building was moved approximately 10–16 m from its original location by the tsunami (Onagawa, 2013). The building was submerged by 0.2 m of seawater at high tide for a short period after the tsunami. It was a four-story steel-frame building with a pile foundation, as shown in Figure 2C. The pile foundation had 8 pile caps with 20 piles having a diameter of 25 cm. Most hollow concrete pipe piles failed at the connection to the pile cap under the foundation unless one pile on the upper right was pulled out. A spiral reinforcing bar and six longitudinal reinforcing bars were observed inside each pile, and the six reinforcing bars had ruptured. This building was constructed from steel frames and ALC walls; thus, building self-weight was less than that of RC buildings. This building was floated, carried away and then overturned by tsunami flow, and most of the piles were broken at their joints with the pile caps (Tokimatsu et al., 2012).

Building D is believed to have been built between 1965 and 1975 (Nikkei BP Company, 2011). It was used as a refrigerated warehouse with a two-story RC structure and a pile foundation, as shown in Figure 2D. The pile foundation had six pile caps with four piles with a diameter of 20 cm in each pile cap. All piles were broken at the pile caps. No spiral reinforcing bar was observed inside the piles, but six longitudinal reinforcing bars were observed inside each pile. This building was floated more than 1 m and moved approximately 7 m, and all piles were ruptured at or near the joints (Tokimatsu et al., 2012). No pile remained connected to the pile caps, suggesting a higher level of shear in the overturning motion than was experienced in the other overturned building with a pile foundation (Fraser et al., 2013). This building was lifted by the hydrostatic buoyancy off of its pile foundation, which did not have tension capacity due to the minimal reinforcing steel (Chock et al., 2013). In addition, it was lifted off its original site and carried over a low wall before being deposited approximately 15 m inland from its original location (Chock et al., 2013).

Building E is believed to have been built in 1980 (Onagawa, 2013). It was a police box on the first floor and a rest area on the second floor (Onagawa, 2013). The building was overturned near its present location and then moved to its present location by the receding wave (Onagawa, 2013). Some damage from floating debris can be observed on the upper part of the building. The building was submerged by 0.3 m of seawater at high tide for a short period after the tsunami. The building was a two-story RC building with a pile foundation, as shown in Figure 2E. The pile foundation had 6 pile caps with 14 piles having a diameter of 25 cm, and most of the piles were pulled out of the ground. No spiral reinforcing bar was observed inside the piles, but six longitudinal reinforcing bars were observed inside each pile. The settled ground near a neighboring building indicates that liquefaction occurred (Tokimatsu et al., 2012).

Based on a video recorded from the rooftop of a building in Onagawa, a thorough analysis of the tsunami inundation was conducted using numerical modeling and measurements (Adriano et al., 2016). Numerical tsunami simulations were performed to reproduce the calculated time series of tsunami flow using the tsunami source model proposed by the Cabinet Office, Government of Japan. The simulations demonstrate that the maximum inundation depth due to the first incoming wave was over 16 m, and more than 500 buildings were washed away by this first wave, which is consistent with the video data (Adriano et al., 2016). The source model was verified with the observed tsunami inundation depth and flow velocity interpreted from the video record including the measured depth of maximum tsunami inundation of 30 points from tsunami watermark and the inundation area measured by field survey and satellite image analysis (Adriano et al., 2016). This study extended this reliable source model of tsunami numerical simulations to reproduce the waveforms of tsunami inundation depth and flow velocity at each overturned building. Then, these waveforms were used to estimate hydrodynamic and buoyancy forces in time series in order to investigate the possible mechanism of building overturning.

The tsunami inundation depth and flow velocity in the time series at each overturned building validated by the interpretation of the video data are shown in Figure 3. The peak inundation depth and peak flow velocity occurred at different times, which could result in the induction of large hydrodynamic force at any time during the striking or receding wave. Figure 3A shows the calculated time series of the tsunami inundation depth and flow velocity in 2014 (Adriano et al., 2014). In this 2014 simulation, the maximum inundation depth was not consistent with the video analysis, although the maximum flow velocity was somewhat consistent. Figure 3B shows the calculated time series of the tsunami inundation depth and flow velocity in 2016, including the crustal deformation (Adriano et al., 2016). In this 2016 simulation, the maximum inundation depth was consistent with the video analysis, whereas the maximum flow velocity was lower than that of the video analysis. However, these numerical tsunami simulations were generated by the tsunami source model, which is typically used to reproduce tsunami propagation for all affected areas in Japan after the 2011 Great East Japan tsunami. This source model is not specific to only the studied area in Onagawa, which makes it different from other reverse models that attempt to be consistent with the video evidence. In this study, sets of inundation depth and flow velocity from both 2014 and 2016 simulations were used to estimate hydrodynamic and buoyancy forces.

The overturning moment is partially the result of hydrodynamic force, which can be calculated from the inundation depth and flow velocity, as shown in Figure 3. This study assumed that these overturned buildings were surrounded by water and had a minimum unbalanced hydrostatic force, i.e., a minimum tsunami load. For each overturned building, hydrodynamic force (F_d) was applied as a uniform load over the depth of tsunami flow (Federal Emergency Management Agency, 2011) as Fd=12ρsCdB(hu2), where ρ_s is the density of salt water with sediment (1,200 kg/m³), C_d is the drag coefficient (2.0), B is the building width in the plane normal to the direction of flow, h is the inundation depth or building height, and u is the flow velocity.

The overturning moment is also partially the result of buoyancy force, which can be calculated from the inundation depth shown in Figure 3. For each overturned building, buoyancy force (F_b) was set equal to the water weight of the residual air space inside it, which also depends on the opening ratio. A residual air space ratio (C_b) of 0 indicates that the entire building was filled with water, whereas a residual air space ratio of 1.0 indicates that no water entered the building, expressed as Fb=ρsCbgBDh, where ρ_s is the density of salt water with sediment (1,200 kg/m³), C_b is the residual air space ratio varying with the relative volume of entrapped air inside the building (0.0–1.0), g is the gravitational acceleration (9.81 m/s²), B is the building width, D is the building depth, and h is the inundation depth or building height.

The resisting moment is also partially the result of building self-weight, which can be calculated from the weight per unit area of the RC and steel buildings. The concrete density is typically 2,400 kg/m³. In this study, the weight per unit area of the RC buildings is assumed to be 14 kPa, whereas the weight per unit area of the steel buildings is assumed to be 8 kPa.

The resisting moment is also partially the result of pile resistance force, which can be calculated from the pile and soil. From the damage observed in the overturned buildings with a pile foundation, two possible cases of pile damage were identified: tension or shear failure at the pile heads and pulling of the pile from the ground.

In general, the evaluation of soil liquefaction during ground shaking and its effects on a pile foundation are highly complex because the seismic ground motion at the site must be considered. In addition, the dissipation of pore-water pressure and lateral ground spreading must be considered to investigate their effects on the soil–pile interaction, which could be caused by the main shock before the tsunami arrival. Therefore, the liquefaction process might be completed before the tsunami arrival and the shear strength of the soil might have recovered by the dissipation of pore-water pressure. However, it may take longer for the soil to regain its shear strength so it might occur within the several tens of minutes between the main shock and tsunami arrival, such as with the repeated occurrence of soil liquefaction in New Zealand within a year due to the September 2010 Canterbury earthquake and the February 2011 Christchurch earthquake.

Because soil liquefaction is an interaction between soil and water, it takes time for the soil and water to slide out. Based on field surveys after two earthquakes in Japan, Mizutani (2008) reported that the soil liquefaction process took considerably longer than the ground shaking process. In the case of the 1964 Niigata earthquake, the sand boiling started after the ground shaking ended, and a large amount of soil boiling was observed for more than 10 to several tens of minutes. In the case of the 1983 Japan Sea earthquake, there was a report of a large sand boil hole with a diameter of 8 m and a depth of 1.5 m that sprayed soil with to a height of 10 m (Japanese Geotechnical Society, Tohoku Branch, 1986), and the outflow of water lasted for more than half a day. Therefore, the direct damage caused by strong ground shaking occurs over several tens of seconds, whereas the indirect damage caused by soil liquefaction requires much more time, and thus is responsible for minimal or no fatalities.

Yeh et al. (2013) suggested that soil liquefaction due to strong ground shaking of the earthquake that had occurred approximately 40 min prior to the tsunami arrival may have further promoted overturning failure. Fraser et al. (2013) also suggested that although any evidence of liquefaction was washed away in the tsunami, it may have contributed by loosening the soil around the piles prior to the overturning motion. Although, the liquefaction process in Onagawa is still not fully understood, it is worth considering the effect of soil liquefaction based on a building design standard because the 2011 tsunami arrived at Onagawa within approximately 40 min of its initiation, when soil liquefaction could not have been negligible. In this study, cases with and without soil liquefaction will be analyzed and compared.

For low-rise regular buildings, equivalent static seismic loads are sufficient to consider base shear force instead of dynamic loads during the ground shaking. The equivalent static seismic loads can be calculated based on a response spectrum analysis using natural period. Based on the AIJ Recommendations for Loads on Buildings (Architectural Institute of Japan, 2004), a simplified method can be used to evaluate the equivalent static seismic loads from the observed response spectrum at the site and an approximation of natural period (T₁) for each overturned building. Therefore, the approximate base shear force (V_B) can be calculated as VB=0.816SagW, T1=(0.02+0.01αh)h, where S_a is the acceleration response at the base of the foundation, g is the gravitational acceleration (9.81 m/s²), W is building self-weight, and α_h is the ratio of steel-frame height to total building height (h).

The base shear force was compared with total shear capacity of the piles in Section “Tension and Shear Capacities” to investigate the potential shear failure of the piles during ground shaking using a safety factor between total shear capacity (∑Q_u) and base shear force (V_B). Table 4 shows the calculation of total shear capacity and base shear force for Buildings B, C, D, and E. The acceleration response (S_a) was obtained from the response spectrum in Ishinomaki city that was provided by NIED, which is the nearest available data to Onagawa [National Research Institute for Earth Science and Disaster Resilience (accessed 2016)]. Based on the safety factor, all piles of Building D failed during the ground shaking due to large base shear force between the pile heads and pile caps, whereas all piles of Buildings C and E were still in a sufficiently good condition to resist tsunami. In addition, some piles of Building B could have failed during the ground shaking because the safety factor is close to 1.00, as shown in Table 4.

The possible overturning mechanism of Buildings A and D was investigated by comparing the overturning moment calculated from hydrodynamic force (F_d) and buoyancy force (F_b) to the resisting moment calculated from building self-weight (W). Building A had a mat foundation, and all piles of Building D failed during the ground shaking, as shown in Figure 5. Figure 6A shows the tsunami inundation depths and flow velocities at Buildings A and D in the 2014 and 2016 simulations. The residual air space ratio (C_b) was estimated based on the condition of larger buoyancy force than building self-weight, as shown in Figure 6B. The residual air space ratios for Buildings A and D were approximately 0.6 and 0.4, respectively. Figure 6C shows the time series of the overturning moment (M_d and M_b) and the resisting moment (M_w).

As shown in Table 5, a comparison of Building A based on the 2014 simulation reveals that the peak overturning moment (M_d + M_b) that occurred at 15:28 was 1.59 times higher than the resisting moment (M_w), whereas a comparison based on the 2016 simulation suggests that the peak overturning moment that occurred at 15:29 was 1.84 times higher than the resisting moment. The peak overturning moment in the 2014 simulation occurred when the inundation depth was 10.60 m and the flow velocity was 2.83 m/s², whereas the maximum depth and velocity of 13.50 m and 4.57 m/s², respectively, occurred at different times, as shown in Figure 6A. The peak overturning moment in the 2016 simulation occurred when the inundation depth was 11.43 m and the flow velocity was 3.48 m/s², whereas the maximum depth of 16.18 m occurred at different times, as shown in Figure 6A. Building A was only 10.5 m tall; thus, it was overturned when the tsunami flow exceeded the top of the building.

All piles of Building D failed during the ground shaking by shear failure between the pile heads and pile caps, in which the shear strength of a pile (Q_u) was 65 kN, as shown in Table 4. As shown in Table 5, a comparison of Building D based on the 2014 simulation reveals that the peak overturning moment (M_d + M_b) that occurred at 15:27 was 1.33 times higher than the resisting moment (M_w), whereas a comparison based on the 2016 simulation suggests that the peak overturning moment that occurred at 15:29 was 1.59 times higher than the resisting moment. The peak overturning moment in the 2014 simulation occurred when the inundation depth was 9.98 m and the flow velocity was 2.02 m/s², whereas the maximum depth and velocity of 13.05 m and 4.07 m/s², respectively, occurred at different times, as shown in Figure 6A. The peak overturning moment in the 2016 simulation occurred when the inundation depth was 10.43 m and the flow velocity was 2.71 m/s², whereas the maximum depth and velocity of 15.93 m and 3.02 m/s², respectively, occurred at different times, as shown in Figure 6A. Building D was 10.5 m tall; thus, it was overturned when the tsunami flow was lower than the top of the building.

Another observation is that buoyancy force could generate an overturning moment larger than hydrodynamic force, particularly for Building D, as shown in Figure 6C. After building overturning occurred at 15:28, these buildings floated and then moved from their original positions such that buoyancy force was immediately greater than building self-weight, as shown in Figure 6B.

The overturning mechanism of Building C was investigated by comparing the overturning moment calculated from hydrodynamic force (F_d) and buoyancy force (F_b) to the resisting moment calculated from building self-weight (W) and pile resistance force (R_TC). Building C had a pile foundation and all of the piles failed by tension failure at the pile heads, except for one pile at the top-right pile cap, as shown in Figure 7. Figure 8A shows the tsunami inundation depth and flow velocity at Building C in the 2014 and 2016 simulations. The residual air space ratio (C_b) was approximately 0.5, as estimated based on the condition of larger buoyancy force than building self-weight, as shown in Figure 8B. Figure 8C shows the time series of the overturning moment (M_d and M_b) and the resisting moment (M_w and M_r).

For Building C, the pile foundation failed during tsunami flow through the observed tension failure between the pile heads and pile caps in which the tensile strength of a pile (F_u) was 307 kN, as shown in Table 4. The effective piles out of the 20 piles were used to calculate the resisting moment (M_r), with 4.0-m moment arm on the pile group equal to the pile length, as shown in Figure 7. As shown in Table 5, a comparison of Building C based on the 2014 simulation reveals that the peak overturning moment (M_d + M_b) that occurred at 15:27 was 1.17 times higher than the resisting moment (M_w + M_r), whereas a comparison based on the 2016 simulation suggests that the peak overturning moment that occurred at 15:30 was 1.24 times higher than the resisting moment. The peak overturning moment in the 2014 simulation occurred when the inundation depth was 8.61 m and the flow velocity was 3.77 m/s², whereas the maximum depth and velocity of 13.18 m and 4.43 m/s², respectively, occurred at different times, as shown in Figure 8A. The peak overturning moment in the 2016 simulation occurred when the inundation depth was 12.14 m and the flow velocity was 1.95 m/s², whereas the maximum depth and velocity of 16.27 m and 2.43 m/s², respectively, occurred at different times, as shown in Figure 8A. Building C was 14.2 m tall; thus, it was overturned when tsunami flow was lower than the top of the building.

Buoyancy force could generate an overturning moment equal to the resisting moment, as shown in Figure 8C. Therefore, although hydrodynamic force could generate a smaller overturning moment than buoyancy force, the additional overturning moment from hydrodynamic force had a significant impact on building overturning.

The overturning mechanisms of Buildings B and E were investigated by comparing the overturning moment calculated from hydrodynamic force (F_d) and buoyancy force (F_b) to the resisting moment calculated from building self-weight (W) and pile resistance force (R_TC). Building B had a pile foundation, and 12 piles were pulled out of the ground, as shown in Figure 9, whereas 20 piles might have failed as a result of base shear force during the ground shaking because the safety factor was close to 1.00, as shown in Table 4. Building E had a pile foundation, and all piles were pulled out of the ground except for one pile that failed in tension, as shown in Figure 9. Figure 10A shows the tsunami inundation depths and flow velocities at Buildings B and E in the 2014 and 2016 simulations. The residual air space ratio (C_b) was estimated based on the condition of larger buoyancy force than building self-weight, as shown in Figure 10B. The residual air space ratios for Buildings B and E were approximately 0.7 and 0.6, respectively. Figure 10C shows the time series of the overturning moment (M_d and M_b) and the resisting moment (M_w and M_r).

For Building B, boring No. 11 at the nearest location and the assumed pile length of 6 m were used to calculate skin friction capacity (Q_s) of a pile, which was 38 kN/m, for a pile diameter of 20 cm, as shown in Table 3. As shown in Table 5, the comparison of Building B based on the 2014 simulation reveals that the peak overturning moment (M_d + M_b) that occurred at 15:27 was 1.61 times higher than the resisting moment (M_w + M_r), whereas the comparison based on the 2016 simulation suggests that the peak overturning moment that occurred at 15:30 was 1.59 times higher than the resisting moment. The peak overturning moment in the 2014 simulation occurred when the inundation depth was 8.71 m and the flow velocity was 3.73 m/s², whereas the maximum depth and velocity of 13.19 m and 4.90 m/s², respectively, occurred at different times, as shown in Figure 10A. The peak overturning moment in the 2016 simulation occurred when the inundation depth was 12.45 m and the flow velocity was 2.10 m/s², as the maximum depth and velocity of 16.02 m and 3.03 m/s², respectively, occurred at different times, as shown in Figure 10A. Building B was 14.0 m tall; thus, it was overturned when tsunami flow was lower than the top of the building.

For Building E, boring No. 5 at the nearest location and the assumed pile length of 6 m were used to calculate skin friction capacity (Q_s) of a pile, which was 64 kN/m, for a pile diameter of 25 cm, as shown in Table 3. As shown in Table 5, the comparison of Building B based on the 2014 simulation in reveals that the peak overturning moment (M_d + M_b) that occurred at 15:26 was 1.47 times higher than the resisting moment (M_w + M_r), whereas the comparison based on the 2016 simulation suggests that the peak overturning moment that occurred at 15:29 was 1.34 times higher than resisting moment. The peak overturning moment in the 2014 simulation occurred when the inundation depth was 7.06 m and the flow velocity was 3.48 m/s², whereas the maximum depth and velocity of 13.63 m and 3.82 m/s², respectively, occurred at different times, as shown in Figure 10A. The peak overturning moment in 2016 occurred when the inundation depth was 10.57 m and the flow velocity was 3.15 m/s², whereas the maximum depth and velocity were 15.88 m and 3.15 m/s², respectively, occurred at different times, as shown in Figure 10A. Building E was 7.0 m tall; thus, it was overturned when tsunami flow exceeded the top of the building.

The overturning ratio (OR) in Table 5 can be calculated from the ratio between the peak overturning moment and the peak resisting moment. For each of 18 soil boring data, skin friction capacity (Q_s) is shown in Table 3 for pile lengths of 4.0, 6.0, and 8.0 m and including the effect of soil liquefaction (Qs′). However, the conventional method to evaluate the effect of soil liquefaction tends to overestimate liquefaction hazards (Chen et al., 2016). These 18 boring data were used to evaluate the potential of overturning for Buildings B and E based on the ORs, which can be classified as no possibility, low possibility, medium possibility, or high possibility.

Table 6 shows the ORs of Buildings B and E in the 2014 and 2016 simulations considering the pile length and soil liquefaction. For all pile lengths, only boring No. 7 can provide safety from building overturning because the OR was less than 1.00 and including the effect of soil liquefaction. On the other hand, boring No. 11 provided the high possibility of building overturning for all pile lengths and including the effect of soil liquefaction. For borings No. 16 and No. 17, safety can be obtained using pile lengths of 6.0 and 8.0 m instead of 4.0 m, as was used in boring No. 3. However, building overturning could occur at boring No. 3 with pile lengths of 6.0 and 8.0 m in the case of soil liquefaction. Nevertheless, increasing the pile length can reduce the OR when neglecting the effect of soil liquefaction. In particular, for boring No. 14, increasing the pile length from 4.0 and 6.0 m to 8.0 m can prevent building overturning by changing the classification from high possibility to no possibility.

The effect of soil liquefaction generally increased the ORs for most of 18 soil boring data. However, the ORs for borings No. 6 and No. 16 were the same with and without soil liquefaction because soil liquefaction could not decrease of skin friction capacity. In the case of soil liquefaction, increasing the pile length could not affect the potential for building overturning in several borings, such as No. 2, No. 4, No. 8, No. 9, No. 11, and No. 12, because skin friction capacity is constant for all pile lengths, as shown in Table 3. In some borings, such as No. 2 and No. 3 with a pile length of 8.0 m, the OR exceeded 1.00 when soil liquefaction occurred because skin friction force was small and decreased significantly, as shown in Table 3.