General Dynamic Collapse Criterion for Elastic–Plastic Structures Under Double Impulse as Substitute of Near-Fault Ground Motion

A dynamic collapse criterion for elastic–plastic structures under near-fault ground motions is derived analytically by approximately transforming near-fault ground motions into double impulse and using an energy balance law. A negative post-yield stiffness is introduced to treat the P-delta effect in the single-degree-of-freedom (SDOF) model. The principal part of fling-step near-fault ground motions is modeled by a one-cycle sine wave and then a double impulse. The double impulse enables the efficient use of the energy approach in the derivation of compact expressions of complicated elastic–plastic responses of structures with the negative post-yield stiffness. In contrast to the previous work (Kojima and Takewaki, 2016b) for the resonant critical case, a general collapse criterion is provided for the velocity amplitude and the frequency of the double impulse. It is significant that no iteration is needed in the derivation of the dynamic collapse criterion except the solution of transcendental equations. It is shown that discussions on several patterns of dynamic collapse behaviors introduced in the previous critical case are useful for deriving a boundary between the collapse and the non-collapse in the plane of the input velocity and the input frequency. The most important point to be remarked is that the critical state (Kojima and Takewaki, 2016b) corresponding to the non-linear resonance does not necessarily provide the minimum input velocity level with respect to arbitrary impulse timing. The validity of the proposed dynamic collapse criterion is examined by the numerical response analysis for structures under double impulses with collapse or non-collapse parameters.

While earthquake and structural engineering made clear its significance in the real world, e.g., safety check of structures and infrastructures, applied mechanics contributed to the theoretical advancement in this field (Herrmann, 1965;Ishida and Morisako, 1985;Maier and Perego, 1992;Araki and Hjelmstad, 2000;Williamson and Hjelmstad, 2001).
It seems that the first theoretical achievement on collapse of structures subjected to earthquake ground motions was made by Jennings and Husid (1968). They dealt with a single-degreeof-freedom (SDOF) system with an elastic-plastic spring and demonstrated that the P-delta effect makes the natural period of the structure longer. Sun et al. (1973) focused on the free vibration of the SDOF system subjected to an initial impact and made clear the stability or collapse boundary. Extension to multidegree-of-freedom (MDOF) systems has also been attempted (Takizawa and Jennings, 1980;Nakajima et al., 1990).
The tangent stiffness has also attracted some researchers in the investigation of dynamic response of elastic-plastic structures in view of instability. If the tangent stiffness goes into the negative range, residual displacements are induced and accelerated. In addition, there were some discussions that a negative eigenvalue of the tangent stiffness matrix is strongly related to either the accumulation of deformation (Uetani and Tagawa, 1998) or the localization of deformation (Maier and Perego, 1992). Dynamic collapse behaviors and responses of actual and realistic frame models have been investigated by some researchers (Ger et al., 1993;Challa and Hall, 1994;Hall, 1998;Sivaselvan et al., 2009). Through such investigations, various effects, such as non-linear geometric effect, non-linear material behavior, and spread of the plastic zone, were incorporated in the numerical methods.
However, it does not seem that a simple dynamic collapse criterion has been proposed even for a rather simple input except the recent work (Kojima and Takewaki, 2016b). In the paper of Kojima and Takewaki (2016b), a simple closed-form dynamic collapse criterion has been proposed for the double impulse as a simplification of the near-fault ground motion by taking full advantage of the energy balance law. They focused on the non-linear resonant situation (Drenick, 1970;Takewaki, 2002;Moustafa et al., 2010) and showed that several collapse patterns exist.
While the inelastic earthquake responses were analyzed for the steady-state response to a harmonic input or the non-stationary response to a simple sinusoidal input in the 1960s−1970s (Caughey, 1960a,b;Iwan, 1961Iwan, , 1965aTakewaki, 2015a,c, 2016a,b) developed a completely different innovative approach to the peak elastic-plastic response using an energy balance law without solving directly the equations of motion. Furthermore, the resonant and overturning phenomena have been investigated from various viewpoints (Chatzis and Smyth, 2012;Makris and Vassiliou, 2013;Casapulla, 2015;Nabeshima et al., 2016;Casapulla and Maione, 2017).
In this paper, a dynamic collapse criterion for elasticplastic structures under near-fault ground motions is derived analytically by approximately transforming near-fault ground motions into a double impulse and using an energy balance law. A negative post-yield stiffness is introduced to treat the P-delta effect in the single-degree-of-freedom (SDOF) model. The principal part of fling-step near-fault ground motions is modeled by a one-cycle sine wave and then a double impulse. The use of the double impulse enables the efficient use of the energy approach in the derivation of explicit expressions of a complicated elastic-plastic response of structures with negative post-yield stiffness. In contrast to the previous work (Kojima and Takewaki, 2016b) for the resonant critical case, a general collapse criterion is provided for the velocity amplitude and the frequency of the double impulse. It is significant that no iteration is needed in the derivation of the dynamic collapse criterion except the solution of transcendental equations. It is shown that discussions on several patterns of dynamic collapse behaviors introduced in the previous critical case are useful for deriving a boundary between the collapse and the non-collapse in the plane of the input velocity and the input frequency. The most important point to be remarked is that the critical state (Kojima and Takewaki, 2016b) corresponding to the nonlinear resonance does not necessarily provide the minimum input velocity level with respect to arbitrary impulse timing. The validity of the proposed dynamic collapse criterion is examined by the numerical response analysis for structures under double impulses with collapse or non-collapse parameters.
There exist two major advantages of the proposed method against the method using time-history response analysis: (1) if the collapse limit figure is prepared, structural designers can judge the state of collapse or non-collapse at once without time-history response analysis and know the safety factor (both for velocity level and input frequency) for the collapse without many timehistory response analyses, and (2) while time-history response analysis of the structural model with negative post-yield stiffness exhibits the response results sensitive to the time increment of the numerical integration, the proposed method does not have such drawback (transcendental equation can be solved stably). Furthermore, since the proposed collapse limit figure is drawn in a normalized form with respect to input velocity level and input frequency, it can be used for various combinations of structural models and input properties.

DOUBLE-IMPULSE INPUT
As explained in the previous papers (Kojima and Takewaki, 2015a,b;Kojima et al., 2015), two types are distinctive in nearfault ground motions. The fault-parallel fling-step input can FIGURE 1 | Transformation of acceleration wavelets into a series of impulses. (A) Fling-step input (blue) and double impulse (red). (B) Forward-directivity input (blue) and triple impulse (red) (Kojima and Takewaki, 2016b).
be expressed by a one-cycle sinusoidal wave (Mavroeidis and Papageorgiou, 2003;Kalkan and Kunnath, 2006) and the faultnormal forward-directivity input can be represented by three sinusoidal wavelets (see Figure 1). The fling step results from the permanent displacement of the ground induced by the fault dislocation and the forward directivity effect can be explained by the relation of the movement of the rupture front with the site. In this paper, a double impulse is used following the references Takewaki, 2015a, 2016a;Kojima et al., 2015). The double impulse enables the derivation of a straightforward expression of the elastic-plastic responses based on an energy approach by taking advantage of the properties of free vibrations. Another advantage of the double impulse is the ease of the derivation of the critical resonant timing of impulses which is not possible for the sinusoidal and other inputs without a repetitive procedure. While most of the conventional methods (Caughey, 1960a,b;Iwan, 1961) use the equivalent linearization techniques for the structural model under the original unchanged input, the method using the double impulse (Kojima and Takewaki, 2015a,b) employs the transformation of the input for the unchanged structural model. This property leads to an advantageous feature that the method using the double impulse is appropriate even for large plastic deformation.
Consider a simplified ground accelerationü g (t) as shown in Figure 1A (Kojima and Takewaki, 2015a) which is expressed bÿ V is the velocity amplitude in both positive and negative directions, and t 0 is the time interval of two impulses. The time derivative is denoted by an over-dot. Figure 1A also illustrates the comparison with the corresponding one-cycle sinusoidal wave. For reference, the velocity and displacement of both inputs are plotted in Figures 1A,B indicates the triple impulse as a substitute of a forward-directivity input. Good agreement can be observed even in the form of velocity and displacement. However, the correspondence in the response should be discussed carefully (see Kojima and Takewaki, 2016a).
The Fourier transform of the accelerationü g (t) of the double impulse can be derived as, Kojima and Takewaki (2016b) treated the case of non-linear resonant input to derive the collapse limit of a bilinear hysteretic SDOF model with negative second slope under the double impulse. The terminology "nonlinear resonant input" in this paper means the case where the non-linear response of the SDOF model attains the maximum with respect to the interval of two impulses of the double impulse. When the restoring-force in the second stiffness range becomes zero, the FIGURE 2 | Regions of collapse and non-collapse and several patterns of collapse limit (patterns of collapse). model collapses. This is because the model cannot sustain the external force at this stage (Kojima and Takewaki, 2016b). Let f and u denote the restoring force of the spring of stiffness k and the displacement of the mass m, respectively. In addition, let f y and d y denote the yielding force and the yield displacement, respectively. The natural period of this SDOF model is denoted by T 1 = 2π/ω 1 (ω 1 = k/m : natural circular frequency). The ratio of the second slope to the initial slope is expressed by α. The negative stiffness depends on the magnitude of elastic stiffness and the effect of the P-effect. When the magnitude of the effect of the P-effect against the elastic stiffness is large, the postyield stiffness is apt to become negative. V y denotes the velocity level at which the SDOF model just attains the yield level after the first impulse. In this non-linear resonant case, the second impulse acts at the point of zero restoring force in the first slope range with a positive slope. They classified the collapse pattern into five patterns, i.e., pattern 1, pattern 2, pattern 3, additional pattern 1, and additional pattern 2. Pattern 1 is the collapse pattern such that the SDOF model collapses after the second impulse without plastic deformation after the first impulse. Pattern 2 is the collapse pattern such that the SDOF model collapses after the second impulse with plastic deformation after the first impulse. Pattern 3 is the collapse pattern such that the SDOF model collapses after the second impulse with plastic deformation after the first impulse and with closed loop after the second impulse. Additional pattern 1 is the collapse pattern such that the SDOF model collapses after the first impulse. Additional pattern 2 is the collapse pattern such that the SDOF model has an elastic limit after the second impulse. CASE 1 indicates the input velocity range such that the SDOF model remains elastic even after the second impulse. CASE 2 expresses the input velocity range such that the SDOF model remains elastic after the first impulse and goes into the plastic range after the second impulse. CASE 3 presents the input velocity range such that the SDOF model goes into the plastic range after the first impulse.

CLASSIFICATION BASED ON INPUT LEVEL OF DOUBLE IMPULSE AND TIMING OF SECOND IMPULSE
In this section, a preparation for the next section to derive the collapse limit is made. Two classifications, one based on the input level of the double impulse and the other based on the timing of the second impulse, are introduced.
In the first classification based on the input level of the double impulse, three cases exist, i.e., CASE-A, B, C.

CASE-A:
The input velocity level satisfies 0 ≤ V/V y ≤ 1.
The parameter 1 − (1/α) is related to the coefficient such that the structure just attains the collapse under only the first impulse with the input velocity level V/V y = 1 − (1/α) (Kojima and Takewaki, 2016b). This classification is different from CASE-1, 2, 3 in the previous section which are defined for resonant critical input. In non-resonant cases, a different kind of classification is necessary.
In the second classification based on the timing of the second impulse, four cases exist, i.e., CASE-I, II, III, and IV.

CASE-I:
The structure does not yield after the first impulse, and the second impulse acts. CASE-II: The structure goes into the plastic range after the first impulse, and the second impulse acts before the structure attains the maximum displacement or before the structure collapses under only the first impulse. CASE-III: The structure goes into the plastic range after the first impulse, and the second impulse acts while the structure exhibits a harmonic free vibration after the attainment of the maximum displacement. CASE-IV: The structure collapses before the action of the second impulse.

(CASE-A)
In CASE-A ( Figure 3A), it is evident that This is because the vibration after the first impulse is a free vibration of an elastic SDOF model.

(CASE-B)
In CASE-B (1 < V/V y < 1 − (1/α)) ( Figure 3A), the ratio of the time between Points O and A to T 1 can be expressed by In this case, the following condition must be satisfied.
The plastic deformation u p1 after the first impulse can be expressed by using the energy balance law Takewaki, 2015a, 2016b) between the point of the first impulse and the maximum deformation point.
The time history from Point A through Point B and the time t AB between Points A and B are derived next. The equation of motion from Point A through Point B is The energy balance law between Points O and A yields From Equation (7) and the initial condition (u(t = 0) = −d y , u(t = 0) = v A ), the displacement and velocity between Points A and B are The ratio of the time between Points A and B to T 1 can be expressed by The time history after Point B and the times between two principal points are derived next. The equation of motion (free vibration) after Point B (maximum displacement point) can be described by The solution of Equation (13) for the initial condition u(t = 0) = −d y − u p1 ,u(t = 0) = 0 and t = 0 at Point B leads to The following results for the ratios of the times between the principal points to T 1 can be obtained.
Based on these results, the ratio of the critical interval of the double impulse to T 1 can be expressed by Consider CASE-C ( Figure 3A). As in CASE-B, the condition (5) is satisfied. The displacement and velocity after yielding (Point A) can be expressed by Equations (10), (11) as in CASE-B. Since the displacement at the collapse (Point D) is −dy{1−(1/α)} and u(t AD ) = −dy{1 − (1/α)} for time t from Point A, the ratio of the time between Point A and Point D to T 1 can be obtained as Furthermore, CASE-III exists in CASE-B and CASE-IV exists in CASE-C. Once such classification of regions based on the timing of the second impulse is conducted, the judgment of collapse or non-collapse is made efficiently by introducing the energy balance law Takewaki, 2015a, 2016b).

DETERMINATION OF COLLAPSE LIMIT INPUT VELOCITY OF DOUBLE IMPULSE WITH ARBITRARY INTERVAL
Consider here several collapse patterns, Collapse patterns 1 ′ -4 ′ . This naming comes from the similarity to the previous formulation for the non-linear resonant case (Kojima and Takewaki, 2016b). Collapse pattern 4 ′ represents a new type.
Collapse Pattern 1 ′ The first collapse pattern is the case where the structure remains elastic after the first impulse and attains the collapse limit after the second impulse with arbitrary timing as shown in Figure 4A.
Let O and A denote the point of the first impulse (origin of the restoring force characteristic) and the point of initial yielding in the negative direction. The interval of two impulses is denoted by t 0 , and the passing time between Points O and A is indicated by t OA . Since the structure does not yield after the first impulse, the following two cases exist.
Consider the respective cases shown in Figure 3.
[0 ≤ V/V y ≤ 1 (CASE-A)] Figure 4A shows the collapse pattern 1 ′ in CASE-A. In this figure, the timing of the second impulse is also indicated. In CASE-A (0 ≤ V/V y ≤ 1), the structure does not collapse for the input 0 < V/V y ≤ 0.5. Therefore, the condition 0.5 < V/V y ≤ 1 is necessary to satisfy the collapse condition.
The displacement and velocity of the mass just before the second impulse can be expressed by When the structure just attains a zero restoring force after the second impulse, the plastic deformation u p2 after the second impulse can be obtained as The energy balance law between the state just after the second impulse and the collapse Point H in Figure 4A can be expressed by m(v * + V) 2 /2 + ku * 2 /2 = (f y d y /2) + f y u p2 + (αku p2 2 /2) (24)  (21), (22), and (23) into Equation (24) leads to the collapse limit input velocity for CASE-A in collapse pattern 1 ′ .

Substitution of Equations
Since 0.5 < V/V y ≤ 1 is necessary, the following condition for α and t 0 must be satisfied.
1/α ≥ 2 cos(2πt 0 /T 1 ) − 1 (26) Figure 4B shows the collapse limit input velocity for α = −0.4 for CASE-A in collapse pattern 1 ′ . It can be observed that the critical case for t 0 /T 1 = 0.5 (Kojima and Takewaki, 2016b) gives the minimum collapse limit input velocity. In Figure 4B, the timing of the second impulse is also indicated.
[1.0 < V/V y (CASE-B, C) and 0 < t 0 ≤ t OA ] After some manipulation, it was found that α ≤ −1 is required in this case to satisfy the collapse condition in collapse pattern 1 ′ . Since a usual case corresponds to the model with α > −1, the detail of analysis is not shown here.

Collapse Pattern 2 ′
The second collapse pattern is the case where the structure exhibits plastic deformation after the first impulse and attains the collapse limit after the second impulse (see Figure 5A). Since the structure exhibits plastic deformation after the first impulse in this case, V/V y > 1 must be satisfied.
Because the second impulse acts after the structure goes into a plastic region under the first impulse, the case is divided into the following two cases, CASE-II and CASE-III.

/α) (CASE-B) and (CASE-II)]
Since the second impulse acts before the attainment of the maximum displacement (Point B) after the yielding under the first impulse, the interval of two impulses has to satisfy The displacement and velocity at time t * just before the action of the second impulse can be obtained from Equations (10), (11).
t OA in Equations (29), (30) can be obtained from Equation (4). The plastic deformation after the first impulse can be expressed by When the maximum displacement after the second impulse just attains a zero restoring force, the plastic deformation after the second impulse can be expressed by The energy balance law between the point just after the second impulse and the point H where the maximum displacement after the second impulse just attains a zero restoring force can be expressed by Substitution of Equations (4), (29)-(32) into Equation (33) provides the collapse input velocity level V/V y for collapse pattern 2 ′ . After some manipulation, it was found that α ≤ −1 is required in this case to satisfy the collapse condition in collapse pattern 2 ′ . Since a usual case corresponds to the model with α > −1, the detail of analysis is not shown here.

[ 1 − (1/α) ≤ V/V y (CASE-C) and (CASE-II)]
Since the second impulse acts before the attainment of the collapse point D with a zero restoring force after the yielding under the first impulse, the following condition must be satisfied from Equations (4), (19).
After some manipulation, it was found that α ≤ −1 is required in this case to satisfy the collapse condition in collapse pattern 2 ′ . Since a usual case corresponds to the model with α > −1, the detail of analysis is not shown here.
[1.0 ≤ V/V y < 1 − (1/α) (CASE-B) and (CASE-III)] Figure 5A shows the collapse pattern 2 ′ in CASE-B. In this figure, the timing of the second impulse is also indicated. In this case, the structure collapses under the second impulse after it goes into a plastic region under the first impulse. Since the second impulse acts after the structure goes into a plastic region and attains the maximum deformation (Point B), Equations (4), (12) require to satisfy the condition In this case, the displacement and velocity of the mass just before the action of the second impulse are described from Equations (14) and (15) as t OA and t AB in Equations (36), (37) can be obtained from Equations (4), (12). The plastic deformation u p1 after the first impulse can be derived by using the energy balance law between the point just after the first impulse and the point of the maximum deformation (Point B). The plastic deformation u p2 after the second impulse can be obtained as Equation (32). The energy balance law between the point just after the second impulse and the point of the maximum deformation u max 2 = d y − u p1 + u p2 after the second impulse (Point H) can be expressed as Since substitution of Equations (4), (6), (12), (32), (36), (37) into Equation (38) provides the transcendental equation, it is difficult to derive a closed-form expression for the input velocity corresponding to the collapse. To determine the input velocity corresponding to the collapse, this transcendental equation can be computed for given α and t 0 . Figure 5B shows the collapse limit input velocity for α = −0.4 for CASE-B in collapse pattern 2 ′ .

Collapse Pattern 3 ′
The third collapse pattern is the case where the structure exhibits plastic deformation after the first impulse and attains the collapse limit with a closed loop after the second impulse (Kojima and Takewaki, 2016b).
Since the structure exhibits plastic deformation after the first impulse in this case, V/V y > 1 must be satisfied. Because the second impulse acts after the structure goes into a plastic region under the first impulse, the case is divided into two cases, CASE-II and CASE-III, as shown in Equation (27). According to the classification shown in Equation (27), the collapse limit velocity corresponding to the collapse pattern 3 ′ is derived. Figure 6A shows the collapse pattern 3 ′ in CASE-B and CASE-II. In this figure, the timing of the second impulse is also indicated. In this case, the structure exhibits a closed loop and collapses under the second impulse after it goes into a plastic region under the first impulse. The second impulse acts before the structure goes into the unloading path naturally and experiences plastic deformation in the positive direction and then in the negative direction.

/α) (CASE-B) and (CASE-II)]
Since the second impulse acts before the structure goes into the unloading path naturally at Point B, the impulse interval t 0 must satisfy Equation (28).
In this case, the displacement and velocity of the mass just before the action of the second impulse are described from Equations (29) and (30). t OA in Equations (29), (30) can be obtained from Equation (4). As shown in Figure 6A, the plastic deformation u p1 after the first impulse can be derived as in Equation (31).
The energy balance law between the point just after the second impulse and the point J in Figure 6A can be expressed by m(v * + V) 2 /2 + k(u * + αu p1 ) 2 /2 = {k(d y − αu p1 ) 2 /2} The plastic deformation u p2 after the second impulse can be obtained from Equation (39). In this case, the condition 0 < u p2 < −(1/α)d y must be satisfied. By solving the quadratic equation, Equation (39), under the condition 0 < u p2 < −(1/α)d y , u p2 can be obtained in closed form.
Another energy balance law between Point J and Point H provides The input velocity corresponding to the collapse can be obtained by solving the quartic equation transformed from Equation (41). Figure 6B shows the collapse limit input velocity for α = −0.4 for CASE-B in collapse pattern 3 ′ .

[ 1 − (1/α) ≤ V/V y (CASE-C) and (CASE-II)]
Since the input velocity level in this case is too large, the solution to satisfy the collapse condition does not exist in this case. Figure 7A shows the collapse pattern 3 ′ in CASE-B and CASE-III. In this figure, the timing of the second impulse is also indicated. In this case, the structure exhibits a closed loop (BIJK) and collapses under the second impulse after it goes into a plastic region under the first impulse. The second impulse acts after the structure goes into the unloading path naturally at Point B. Since the second impulse acts after the structure goes into an unloading path naturally at Point B, the impulse interval t 0 must satisfy Equation (35). The displacement and velocity just before the action of the second impulse can be obtained from Equations (36), (37). t OA and t AB in Equations (36), (37) can be obtained from Equations (4), (12). The plastic deformation u p1 after the first impulse can be derived as Equation (6) by using the energy balance law between the point just after the first impulse and the point of the maximum deformation (Point B) in Figure 7A.

/α) (CASE-B) and (CASE-III)]
The energy balance law between the point just after the second impulse and the point of the maximum deformation u max 2 after the second impulse (Point J in Figure 7A) can be expressed as By solving the quadratic equation, Equation (42), under the condition 0<u p2 <−(1/α)d y , u p2 can be obtained in closed form. The collapse limit level in this pattern can be obtained by solving the quartic equation derived by substituting Equation (43) into Equation (41). Figure 7B shows the collapse limit input velocity for α = −0.4 for CASE-B in collapse pattern 3 ′ .

Collapse Pattern 4 ′
The fourth collapse pattern is the case where the structure exhibits plastic deformation after the first impulse and attains the collapse limit after experiencing unloading (positive direction) and reloading-reyielding (negative direction) for the second impulse.
Since the structure exhibits plastic deformation after the first impulse in this case, V/V y > 1 must be satisfied. Because the second impulse acts after the structure goes into a plastic region under the first impulse, the case is divided into two cases, CASE-II and CASE-III, as shown in Equation (27). According to the classification shown in Equation (27), the collapse limit velocity corresponding to the collapse pattern 4 ′ is derived.
[1 < V/V y < 1 − (1/α) (CASE-B) and (CASE-II)] Figure 8A shows the collapse pattern 4 ′ for CASE-B and CASE-II. In this figure, the timing of the second impulse is also indicated. In this case, the second impulse acts before the structure goes into an unloading path naturally. The structure does not experience plastic deformation in the positive direction. Since the second impulse acts before the structure goes into an unloading path naturally at Point B, the impulse interval t 0 must satisfy Equation (28).
In this case, the displacement and velocity of the mass just before the action of the second impulse are expressed by Equations (29) and (30). t OA in Equations (29), (30) can be obtained from Equation (4). As shown in Figure 8A, the plastic deformation u p1 after the first impulse can be derived as in Equation (31). Since the structure does not go into a plastic region just after the second impulse, the following relation must hold. m(v * + V) 2 /2 + k{u * + αu p1 } 2 /2 ≤ {k(d y − αu p1 ) 2 /2} (44) The energy balance law between the point just after the second impulse and Point H in Figure 8A can be expressed by [ 1 − (1/α) ≤ V/V y (CASE-C) and (CASE-II)] Figure 8B shows the collapse pattern 4 ′ for CASE-C and CASE-II. In this figure, the timing of the second impulse is also indicated. In this case, the second impulse acts before the structure attains the collapse state, Point D. The structure does not experience plastic deformation in the positive direction.
Since the second impulse acts before the structure attains the collapse state, Point D, the impulse interval t 0 must satisfy Equation (34).
In this case, the displacement and velocity of the mass just before the action of the second impulse are described by Equations (29) and (30). t OA in Equations (29), (30) can be obtained from Equation (4). As shown in Figure 8B, the plastic deformation u p1 after the first impulse can be derived as in Equation (31). Since the structure does not go into a plastic region just after the action of the second impulse, Equation (44) must be satisfied. The energy balance law between the point just after the FIGURE 11 | Second impulse timing t 0 /T 1 -input velocity relation for collapse and non-collapse states (α = −0.4).
FIGURE 12 | Correspondence between arbitrary timing and critical timing of second impulse in collapse velocity level (α = −0.4).
Frontiers in Built Environment | www.frontiersin.org  second impulse and the point H in Figure 8A can be expressed as Equation (45). Figure 9 shows the collapse limit input velocity for α = −0.4 for Case-B, CASE-II and CASE-C, CASE-II in collapse pattern 4 ′ .
[1 ≤ V/V y < 1 − (1/α) (CASE-B) and (CASE-III)] Figure 10A shows the collapse pattern 4 ′ for CASE-B and CASE-III. In this figure, the timing of the second impulse is also indicated. In this case, the second impulse acts after the structure goes into a plastic range and attains the maximum deformation, Point B. The structure does not experience plastic deformation in the positive direction. Since the second impulse acts after the structure attains the maximum deformation, Point B, the impulse interval t 0 must satisfy Equation (35).
When the structure goes into a plastic region after the first impulse and the second impulse acts after the structure attains the maximum deformation, Point B, the displacement and velocity of the mass just before the action of the second impulse are described by Equations (36) and (37). t OA and t AB in Equations (36), (37) can be obtained by Equations (4), (12). As shown in Figure 10A, the plastic deformation u p1 after the first impulse can be derived as in Equation (6) by using the balance law between the point just after the first impulse and the point B of the maximum deformation.
Since the structure does not go into a plastic region just after the action of the second impulse, Equation (46) must be satisfied.
The energy balance law between the point just after the second impulse and the point H in Figure 10A can be expressed as Since substitution of Equations (4), (6), (12), (32), (36), (37) into Equation (47) provides the transcendental equation, it is difficult to derive a closed-form expression for the input velocity corresponding to the collapse. To determine the input velocity corresponding to the collapse, this transcendental equation can be computed for given α and t 0 . Figure 10B shows the second impulse timing t 0 /T 1input velocity relation for Case-B and CASE-III in collapse pattern 4 ′ .

COLLAPSE LIMIT INPUT VELOCITY OF DOUBLE IMPULSE WITH ARBITRARY INTERVAL FOR SDOF SYSTEM WITH VARIOUS NEGATIVE POST-YIELD SLOPES
Based on the collapse patterns explained above, a limit curve on the second impulse timing t 0 /T 1 -input velocity V/V y relation for the collapse and non-collapse states can be proposed. As an example, Figure 11 shows the second impulse timing t 0 /T 1input velocity relation for the collapse and non-collapse states for α = −0.4. It should be remarked that the present SDOF model is an undamped model, and the states of t 0 /T 1 =0.5 and t 0 /T 1 =1.5 provide the same collapse limit. It can be observed that an isolated region of the collapse state exists around the level of t 0 /T 1 =0.5 (also 1.5) and the level of V/V y =1. The most important point to be remarked is that the critical state (Kojima and Takewaki, 2016b) corresponding to the non-linear resonance does not necessarily provide the minimum input velocity level with respect to arbitrary impulse timing. Figure 12 presents the correspondence between arbitrary timing and the critical timing (Kojima and Takewaki, 2016b) of the second impulse in the collapse velocity level (α = −0.4). It can be confirmed that the states of t 0 /T 1 =0.5 and t 0 /T 1 =1.5 certainly correspond to the critical state in the reference (Kojima and Takewaki, 2016b). Figure 13 indicates the collapse limit input velocity of the double impulse with arbitrary interval for the SDOF system with various negative post-yield slopes (α = −0.2, −1/3, −0.8). The correspondence between the arbitrary timing and critical timing (Kojima and Takewaki, 2016b) of the second impulse in the collapse velocity level is also shown again. It can be found that, as the parameter α changes, different phases of the limit curve in the second impulse timing t 0 /T 1input velocity relation for collapse and non-collapse states appear and α = −1/3 gives the boundary of the change of phases. When α is larger than −1/3, the non-linear resonance does not provide the minimum input level corresponding to collapse. Figure 14A shows the verification of the proposed collapse limit by time-history response analysis to double impulses with various input velocities and impulse timings (α = −0.2, −0.4). Many combinations of the impulse timing and the input velocity level in the collapse and non-collapse states were selected in the case of α = −0.2, −0.4. For strict verification, many combinations of the impulse timing and the input velocity level were chosen near the boundary of the proposed limit curve. Figure 14B presents the displacement time histories of the model with α = −0.4 to five inputs (1), (2), (5), (6), (8) shown in Figure 14A. It can be observed that strict classification into the collapse state and the non-collapse state has been made with the proposed limit curve. Figure 15 presents the 9 restoring-force characteristics corresponding to various combinations of the impulse timing and the input velocity level shown in Figure 14 (blue marks indicate the double impulse timings and red mark shows the collapse state). The cases (1), (2), (5), (6), (8) exhibit collapse behaviors. The response behaviors can be well-understood from these figures.

CONCLUSIONS
A dynamic collapse criterion for elastic-plastic structures under double impulse as a substitute of a near-fault ground motion has been derived. The conclusions may be summarized as follows: (1) The use of the double impulse enables the efficient use of the energy approach in the derivation of explicit expressions of a complicated elastic-plastic response of structures with the P-delta effect.
(2) In contrast to the previous work (Kojima and Takewaki, 2016b) for the resonant critical case, a general collapse criterion is provided for the velocity amplitude and the frequency of the double impulse. It is significant that no iteration is needed in the derivation of the dynamic collapse criterion. (3) Discussions on several patterns of dynamic collapse behaviors introduced in the previous critical case are useful for deriving a boundary between the collapse and the non-collapse in the plane of the input velocity and the input frequency. (4) The most important point to be remarked is that the critical state (Kojima and Takewaki, 2016b) corresponding to the non-linear resonance does not necessarily provide the minimum input velocity level with respect to arbitrary impulse timing.
(5) The validity of the proposed collapse criterion has been investigated by the numerical response analysis for structures under double impulses with collapse or noncollapse parameters. It has been confirmed that the proposed criterion has a reasonable accuracy.
The present paper dealt with an undamped system. This is because, if a damped system is treated, the formulation is too complicated even for the case of critical input (Saotome et al., 2019). The discussion on the damped system may be the future work.

DATA AVAILABILITY STATEMENT
All datasets generated for this study are included in the article/supplementary material.