Optimum Strength Distribution for Structures with Metallic Dampers Subjected to Seismic Loading

Abstract

A key aspect of the seismic design of structures is the distribution of the lateral strength, because it governs the distribution of the cumulative plastic strain energy (i.e., the damage) among the stories. The lateral shear strength of a story i is commonly normalized by the upward weight of the building and expressed by a shear force coefficient α_i. The cumulative plastic strain energy in a given story i can be normalized by the product of its lateral strength and yield displacement, and expressed by a plastic deformation ratio η_i. The distribution α_i/α₁ that makes η_i equal in all stories is called the optimum yield-shear force distribution. It constitutes a major aim of design; a second aim is to achieve similar ductility demand in all stories. This paper proposes a new approach for deriving the optimum yield-shear force coefficient distribution of structures without underground stories and equipped with metallic dampers. It is shown, both numerically and experimentally, that structures designed with the proposed distribution fulfil the expected response in terms of both damage distribution and inter-story drift demand. Moreover, a comparison with other distributions described in the literature serves to underscore the advantages of the proposed approach.

1. Introduction

Conventional seismic design allows for damage to a structure under rare or very rare seismic events for economic reasons. The philosophy is to design the structural elements (e.g., beams and columns in frames) so that they can dissipate most of the energy introduced by the earthquake through plastic deformations, without leading to the collapse of the structure. To the present day, the seismic design of structures is made with the so-called force-based approach, which characterizes the loading effect of earthquakes on structures in terms of forces. The force-based approach implements the above philosophy through the use of ductility factors and the assumption of the “equal-displacement” rule. The origin of this approach can be found in the studies carried out for single degree of freedom (SDOF) systems by Veletsos and Newmark [1]. The equal displacement rule assumes that the maximum displacement of a SDOF elastoplastic system and that of the counterpart elastic system under a given ground motion are the same. Based on this assumption, the design lateral strength of the structure is reduced by a ductility-based factor μ (called q in Eurocode 8 [2]) with respect to the value that would be required for the structure to remain elastic. However, many difficulties arise when this approach is extended to multi-story structures. One resides in how to predict the distribution of the cumulative plastic energy dissipation demand E_hi among stories. Controlling this distribution is a key aspect of earthquake resistant design [3]. Force-based methods resolve the inability to deal explicitly with E_hi by prescribing rules aimed at avoiding/alleviating the concentration of plastic strain energy demands in specific stories. These rules govern both the global (the whole structure) and the local level (the individual structural elements). Global level requirements include: (i) a minimum ratio between the sum of the flexural strengths of the columns, ΣM_Rc, and the flexural strengths of the beams, ΣM_Rb, that frame at the same joint; (ii) avoidance of abrupt variations in lateral stiffness and mass distributions, and (iii) ensuring that, at all end sections in a building where plastic hinges may form, the ratio between the maximum and minimum values of the parameter ρ_i = D_i/C_i is between 2 and 3 [2]. Here, D_i is the demand obtained from a linear elastic analysis (e.g., the bending moment in moment frames), and C_i the capacity (e.g., the moment resistance in moment frames). Local level requirements include the capacity design of the sections. Yet, for economic reasons, the required minimum value of the ratio ΣM_Rc/ΣM_Rb in codes (about 1.3) is far below the one that would guarantee the formation of a full strong-column/weak-beam mechanism (≥4) [4]. Hence, damage concentrations in given stories can occur to some extent.

Contrary to the presumption of the force-based approach that an earthquake exerts forces on the structure, in reality it imparts seismic energy [5]. Characterizing the loading effect of earthquakes in terms of energy instead of forces provides a more rational ground for seismic design and constitutes the basis of the so-called energy-based approach. This approach, first proposed by Housner [6] and later developed by Akiyama [7], Bertero [3] and others, has been gaining attention in the last two decades. It was implemented in 2009 in the Japanese Building Code [8]. Energy-based concepts are crucial also to analyse and overview strategies for designing buildings with energy dissipation devices, as shown for example in a recent review on different strategies for the optimal placement of viscous dampers in buildings structures [9]. The fact that the energy-based approach addresses the prediction of E_hi directly constitutes the core of the methodology. For convenience, E_hi can be normalized by the product of the lateral strength Q_yi and yield displacement δ_yi, of the story and expressed by the parameter η_i = E_hi/(Q_yiδ_yi). The ratio η_i has been widely used in the past as an index of damage. Akiyama [7] showed that the distribution of η_i among stories depends on the extent to which the so-called yield-shear force coefficient of the story, α_i, deviates from an “optimum” value α_opt,i, Here, α_i is defined as Q_yi normalized by the weight above the story, i.e., ${\alpha }_i=Q_{yi}/{{\displaystyle \sum }}_{j=i}^N{m}_{j}g$, where m_j is the story mass, g the acceleration of gravity and N the total number of stories. α_opt,i is defined as the value of α_i that makes η_i equal in all stories (i.e., uniform distribution of damage). The distribution of α_opt,i, i.e., ${\overline{\alpha }}_i={\alpha }_{opt,i}/{\alpha }_1$, is referred to as the “optimum yield-shear force coefficient distribution”, and is a key aspect of seismic design methodology based on the energy approach. While several general expressions of ${\overline{\alpha }}_i$ have been proposed in the past for conventional structures [7,8,10,11], optimum distributions specifically for structures with energy dissipation systems are very scarce [12], in particular for those using metallic energy dissipation devices (dampers). The structures with metallic dampers constitute an innovative seismic design strategy whose use has increased exponentially in recent decades. Structures with metallic dampers solve one of the main drawbacks of conventional structures—the occurrence of plastic deformations in beams and columns of the frame—and concentrate the plastic-strain energy dissipation demands in specific elements (dampers) specially designed for that purpose and easily replaceable.

This paper proposes a new expression for the optimum yield-shear force coefficient distribution of structures without underground stories and equipped with metallic dampers, for application in the framework of the energy-based seismic design approach. In contrast to previous approaches that resort to regression analyses of the responses obtained through iterative trial-and-error calculations of structures subjected to a set of ground motions, the optimum yield-shear force distribution proposed in this study involves applying modal analysis formulation and two basic assumptions: (i) damage basically spreads out evenly among the stories regardless of the level of plastic deformation; and (ii) the ductility demand is approximately the same in all stories. The validity of these assumptions for structures designed with the optimum distribution has been demonstrated in past studies [7]. The optimum yield-shear force distribution proposed in this study applies a novel approach in comparison with the optimum distribution developed by the authors in previous studies [12]. In Benavent-Climent [12] the building was represented by an equivalent shear strut (continuous model with distributed mass and stiffness) fixed at the base and the optimum strength distribution was obtained assuming that it coincides with the maximum shear-force distribution in the equivalent elastic undamped shear strut. The equation proposed in [12] approximated the “exact” curve that is obtained by solving the partial differential equation of the elastic undamped distributed parameter system. The new optimum distribution proposed in this paper is obtained representing the building by a discrete model with the mass lumped at several points and the lateral stiffness represented by discrete springs, under the two assumptions motioned above. The new optimum distribution applies modal analysis formulation without resorting to approximations. The application of the proposed optimum distribution to buildings with underground stories, for which the soil-structure interaction could be significant, is beyond the scope of this study. The proposed optimum distribution can be used in buildings with uneven height of the story.

2. Background

2.1. Idealization of a Structure with Metallic Dampers

Structures with metallic dampers consist of two parts working in parallel. The “main structure” is entrusted with sustaining the gravity loads while the structure moves laterally, remaining basically elastic. The other part is the “energy dissipation system”, entrusted with dissipating most of the energy input by the earthquake. The latter comprises the dampers and the auxiliary elements needed to connect them with the “main structure”. Figure 1 shows in bold lines the idealized curve that represents the i-th story shear force versus inter-story drift, Q_i-δ_i, of a structure with metallic dampers subjected to monotonic (Figure 1a) and cyclic (Figure 1b) loading. This curve is the sum of two curves. One represents the main structure, _fQ_i-δ_i, and is characterized by the elastic stiffness _fk_i. The other represents the energy dissipation system, _sQ_i-δ_i, and is defined by the initial stiffness _sk_i, the yield strength, _sQ_y,i, and the yield displacement _sδ_y,i. The stiffness ratio K_i is defined by K_i = _sk_i/_fk_i. The “energy dissipation system” is typically made stiffer laterally than the “main structure”. Accordingly, hereafter, the “energy dissipation system” will be designated with the subindex “s” (stiff part) and the “main structure” with the subindex “f” (flexible part). The flexible part is assumed elastic and the stiff part is assumed to exhibit elastic-perfectly plastic restoring force characteristics.

The main structure must have a large enough elastic deformation capacity to achieve the maximum lateral displacement expected under seismic loads. As for the strength, past research [7] recommended the following value for the ratio r_q,i between the averaged maximum shear force attained by the main structure in the positive and negative domains,${}_f{\overline{Q}}_{max,i}$(=$[{}_f{Q}_{max,i}+{}_f{Q}_{min,i}$]$/2),$ and the yield-shear force of the energy dissipation system, _sQ_y,i at each story level:

$$r_{q,i}=\frac{{}_f{\overline{Q}}_{max,i}}{{}_f{Q}_{y,i}}\ge 0.8$$

(1)

2.2. Energy-Based Design Framework

The dynamic response of an N-degree of freedom system subjected to a seismic ground motion is governed by the following force-equilibrium equation:

$$\mathbf{M}\ddot{\mathbf{u}}\left(t\right)+\mathbf{C}\dot{\mathbf{u}}\left(t\right)+\mathbf{F}\left(t\right)=-\mathbf{M}r{\ddot{u}}_g\left(t\right)$$

(2)

where M and C are the mass and damping matrices, F(t) is the restoring force vector, r is the directivity ground motion vector, ü_g(t) is the ground acceleration, u is the relative displacement vector respect to the foundation ground and t the time. Pre-multiplying Equation (2) by $d\mathbf{u}{\left(t\right)}^{\mathrm{T}}=\dot{\mathbf{u}}{\left(t\right)}^{\mathrm{T}}dt$ and integrating over the duration of the earthquake, t = t_d, the energy balance equation is obtained:

$${{\displaystyle \int }}_0^{t_d}\dot{\mathbf{u}}{\left(t\right)}^{\mathrm{T}}M\ddot{\mathrm{u}}\left(t\right)dt+{{\displaystyle \int }}_0^{t_d}\dot{\mathbf{u}}{\left(t\right)}^{\mathrm{T}}C\dot{\mathbf{u}}\left(t\right)dt+{{\displaystyle \int }}_0^{t_d}\dot{\mathbf{u}}{\left(t\right)}^{\mathrm{T}}Fdt=-{{\displaystyle \int }}_0^{t_d}\dot{\mathbf{u}}{\left(t\right)}^{\mathrm{T}}\mathbf{M}r{\ddot{u}}_g\left(t\right)dt$$

(3)

where E_k is the kinetic energy, E_ξ is the damping energy, E_a is the absorbed energy and E_I is the relative input energy, (hereafter input energy). Further, E_I can also be expressed as an equivalent velocity, V_E, defined by $V_E=\sqrt{2E_I/M}$, where M is the total mass of the structure, and Equation (3) can be rewritten as:

$$E_k+E_{\xi }+E_a=E_I=\frac{1}{2}M{V}_E^2$$

(4)

where E_a is the sum of the elastic strain energy, E_s, and the energy dissipated by plastic deformations or hysteretic energy, E_h, i.e., E_a = E_s + E_h. By definition, the sum of E_k and E_s is the so-called vibrational elastic energy, E_e (=E_k + E_s). Housner [6] defined the sum of E_e and E_h as the energy that contributes to the damage E_D (=E_e + E_h). E_D can also be expressed in the form of an equivalent velocity, V_D, by $V_D=\sqrt{2E_D/M}$ and thus Equation (4) can be rearranged as follows:

$$E_e+E_h=E_I-E_{\xi }=E_D=\frac{1}{2}M{V}_D^2$$

(5)

Furthermore, several empirical expressions have been proposed in the literature to estimate V_D from V_E [7,13,14,15,16,17,18]. In this study, the one proposed by Akiyama [7] is used:

$$\frac{V_D}{V_E}=\frac{1}{1+3\xi +1.2\sqrt{\xi }}$$

(6)

where ξ is the inherent fraction of damping of the structure. For elastic MDOF systems, F(t) can be expressed by F(t) = Ku(t), where K is the elastic stiffness matrix. The damping matrix C can be in general non-proportional; yet, even in this case it can be approximated by a damping matrix that can be diagonalized by the natural vibration modes. This approximated damping matrix is obtained ignoring the off-diagonal coupling coefficients of the modal coordinate damping matrix. Then, the response vector u can be obtained by superimposing the contribution of each vibrational mode, i.e., $u={{\displaystyle \sum }}_{n=1}^N{\phi }_n{x}_n\left(t\right)$, where x_n(t) is the normal coordinate corresponding to mode n and φ_n the corresponding mode vector. Using a damping matrix C that can be diagonalized by the undamped vibration modes, taking into account the orthogonality properties, and pre-multiplying Equation (2) by φ_n^T, gives the following set of N-uncoupled equations in normal coordinates:

$$m_n{\ddot{x}}_n\left(t\right)+c{\dot{x}}_n\left(t\right)+k_n{x}_n\left(t\right)=-m_n{\mathsf{\Gamma }}_n{\ddot{u}}_g\left(t\right)$$

(7)

where m_n = φ_n^TMφ_n, c_n = φ_n^TCφ_n and k_n = φ_n^TKφ_n are the n-th generalized mass, damping and stiffness, respectively, and Γ_n = φ_n^TMr/m_n is the n-th modal participation factor. The displacements x_n(t) obtained from Equation (7) are Γ_n times larger than those of an equivalent SDOF system, x_SDOF,n(t), with mass, damping and stiffness given by m_n, c_n and k_n, respectively, subjected to the ground acceleration ü_g. That is, x_n(t) = Γ_nx_SDOF,n(t). Then, multiplying both sides of the Equation (7) by $d{x}_n={\dot{x}}_{n}dt$ and integrating between t = 0 and t ≤ t_d, the energy balance equation corresponding to the uncoupled n-th vibration mode is obtained as follows:

$${{\displaystyle \int }}_0^t{m}_n{\ddot{x}}_n\left(t\right){\dot{x}}_n\left(t\right)dt+{{\displaystyle \int }}_0^t{c}_n{\dot{x}}_n\left(t\right){\dot{x}}_n\left(t\right)dt+{{\displaystyle \int }}_0^t{k}_n{x}_n\left(t\right){\dot{x}}_n\left(t\right)dt=-{{\displaystyle \int }}_0^t{m}_n{\dot{x}}_n\left(t\right){\mathsf{\Gamma }}_n{\ddot{u}}_g\left(t\right)dt$$

(8)

The first three terms of the left-hand side are the kinetic energy E_k,n, the damping energy E_ξ,n and the absorbed energy E_a,n corresponding to vibration mode n, respectively. It is worth noting that, given the linear nature of Equation (8), the third term on the left-hand side should be, strictly speaking, denoted as the elastic strain energy, which coincides with the absorbed energy in this case. The term on the right-hand side is the input energy corresponding to vibration mode n, E_I,n. Each term of Equation (8) is Γ_n² times the corresponding value in the n-th equivalent SDOF system; in particular, $E_{an}={\mathsf{\Gamma }}_n^2{E}_{an}^{SDOF}$, where $E_{an}^{SDOF}$ is the absorbed energy of the n-th equivalent SDOF system. Moreover, substituting $u={{\displaystyle \sum }}_{n=1}^N{\phi }_n{x}_n\left(t\right)$ in Equation (3) and taking into account that M, K and C are orthogonal, the energy balance equation of the elastic MDOF systems at t ≤ t_d is:

$${\displaystyle \sum }_{n=1}^N\left[{{\displaystyle \int }}_0^t{m}_n{\ddot{x}}_n\left(t\right){\dot{x}}_n\left(t\right)dt\right]+{\displaystyle \sum }_{n=1}^N\left[{{\displaystyle \int }}_0^t{c}_n{\dot{x}}_n\left(t\right){\dot{x}}_n\left(t\right)dt\right]+{\displaystyle \sum }_{n=1}^N\left[{{\displaystyle \int }}_0^t{k}_n{x}_n\left(t\right){\dot{x}}_n\left(t\right)dt\right]={\displaystyle \sum }_{n=1}^N\left[-{{\displaystyle \int }}_0^t{m}_n{\dot{x}}_n\left(t\right){\mathsf{\Gamma }}_n{\ddot{u}}_g\left(t\right)dt\right]$$

(9)

From Equations (3), (8), and (9), it can be observed that E_k, E_ξ, E_a and E_I in the elastic MDOF systems are obtained summing up the contribution of all n-th vibration mode energies, E_kn, E_ξn, E_an and E_In. In particular, $E_a={{\displaystyle \sum }}_{n=1}^N{E}_{an}$ can also be expressed as follows:

$$E_a={\displaystyle \sum }_{n=1}^N{\mathsf{\Gamma }}_n^2{E}_{an}^{SDOF}$$

(10)

Similar expressions can also be obtained for E_k, E_ξ and E_I. Equation (10) shows that the absorbed energy of an elastic MDOF system can be obtained as the superposition of the absorbed energy of N equivalent SDOF systems, each one associated with a vibration mode n, subjected to the ground acceleration ü_g and scaled by ${\mathsf{\Gamma }}_n^2$. Chou and Uang [19] proposed applying the superposition of E_a,n likewise for structures that undergo plastic deformations to estimate the total absorbed energy at t = t_d. Further, Akiyama [7] showed that as the structure enters the nonlinear range, E_e (=E_k + E_s) becomes negligible compared with E_h at t ≥ t_d and can be disregarded, i.e., E_h = E_a. Therefore, the following Equation (10), E_h can be estimated by:

$$E_h={\displaystyle \sum }_{n=1}^N{E}_{hn}={\displaystyle \sum }_{n=1}^N{\mathsf{\Gamma }}_n^2{E}_{hn}^{SDOF}$$

(11)

where $E_{hn}^{SDOF}$ is the hysteretic energy dissipated by the n-th equivalent SDOF system. If the structure is classically damped, it suffices to add the contribution of the first r vibrational modes that mobilize 90% of the total mass of the structure. In the more general case that the structure is non-classically damped, it has been shown in recent research [20] that the modal mass ratio alone is not sufficient to ascertain the accuracy of a truncated model. A measure that is related to the damping matrix, called modal dissipation ratio of each mode in a complex modal analysis framework, should necessary be considered when constructing a reduced-order model.

2.3. Hysteretic Energy Accumulated under Cyclic and under Monotonic Loads

Under the cyclic reversals caused by the ground motion, each story i dissipates a portion of hysteretic energy E_h,i of the total hysteretic energy $E_h(={{\displaystyle \sum }}_{j=1}^N{E}_{h,j})$ dissipated by the overall structure. The quotient ψ_I = E_h,i/E_h is defined hereafter as the i-th story hysteretic energy ratio. After the ground motion fades away, i.e., at t ≥ t_d, E_h = E_a because the elastic strain energy is totally recovered and E_e = 0, so that ψ_i = E_a,i/E_a. A modal hysteretic energy ratio ψ_n,i can also be defined using the hysteretic energy dissipated at the i-th story in the mode n, E_hn,i, and the total hysteretic energy dissipated in mode n, $E_{hn}(={{\displaystyle \sum }}_{j=1}^N{E}_{hn,j})$, i.e., ψ_n,i = E_hn,i/E_hn, and the remark made above for ψ_i at t ≥ t_d is also applicable to ψ_n,i. Similarly, if the structure is subjected to monotonically applied lateral loads following the n-th vibration mode instead of cyclic loads, the ratio between the hysteretic energy dissipated at the i-th story, E_hmn,i, and the total dissipated energy, $E_{hmn}(={{\displaystyle \sum }}_{j=1}^N{E}_{hmn,j})$, define a new ratio ψ_mn,i as follows:

$${\psi }_{mn,i}=\frac{E_{hmn,i}}{E_{hmn}}=\frac{E_{hmn,i}}{{{\displaystyle \sum }}_{j=1}^N{E}_{hmn,j}}$$

(12)

Throughout numerical simulations conducted on MDOF systems subjected to several ground motions, Chou and Uang [19] showed that ψ_n,i and ψ_mn,i have similar values. Therefore, E_hn,i can be estimated by $E_{hn,i}={\psi }_{mn,i}{E}_{hn}$. Recalling that $E_{hn}={\mathsf{\Gamma }}_n^2{E}_{hn}^{SDOF}$ and $E_{h,i}={{\displaystyle \sum }}_{n=1}^N{E}_{hn,i}$ it follows that:

$$E_{h,i}={\displaystyle \sum }_{n=1}^N{\psi }_{mn,i}{\mathsf{\Gamma }}_n^2{E}_{hn}^{SDOF}$$

(13)

In a general MDOF system subjected to a given ground motion characterized by its hysteretic energy spectrum E_h^SDOF-T, the hysteretic energy accumulated in the i-th story E_h,i can be estimated through Equation (13), first, calculating ψ_mn,i from a pushover analysis with the n-th vibrational mode loading pattern and, second, taking as E_hn^SDOF the ordinate of the E_h^SDOF-T spectrum at period T_n.

3. Proposal of New Optimum Yield-Shear Force Coefficient Distribution

The main purpose of installing dampers is to avoid damage in the main structure and concentrate the plastic deformations in the energy dissipation system. Accordingly, in the following discussion the main structure is assumed to remain elastic and the cumulative plastic strain energy ratioη_i for this type of structure is defined as η_i = _sE_h,i/(_sQ_y,i × _sδ_y,i), where _sE_h,i denotes the hysteretic energy dissipated by the energy dissipation devices (dampers) installed at the i-th story. For convenience, noting that ${}_s{Q}_{y,i}={{\displaystyle \sum }}_{j=i}^N\left(m_{j}g\right){}_s\alpha {}_i$, ${}_s\delta {}_{y,i}={}_{s}Q{}_{y,i}/{}_{s}k{}_i={{\displaystyle \sum }}_{j=i}^N\left(m_{j}g\right){}_s\alpha {}_i/\left(K_i{}_{f}k{}_i\right)$ and using Equation (13), the plastic strain energy ration η_i can be rewritten as follows:

$${\eta }_i=\frac{K_i{}_{f}k{}_i{{\displaystyle \sum }}_{n=1}^r\left({\psi }_{mn,i}{\mathsf{\Gamma }}_n^2{E}_{hn}^{SDOF}\right)}{{\left[{{\displaystyle \sum }}_{j=i}^N\left(m_{j}g\right)\right]}^2{}_s\alpha {{}_i}^2}$$

(14)

where the sum in the numerator is extended to r vibration modes. For a given ground motion, the _sα_i’s that make η_i equal in all stories, i.e., η_i = η, are defined here as the optimum yield-shear force coefficients of the energy dissipation system, _sα_opt,i. The corresponding distribution, _sα_opt,i/_sα₁, is referred to hereafter as optimum yield shear force distribution ${}_s\overline{\alpha }{}_i$ (=_sα_opt,i/_sα₁), where _sα₁ is the yield-shear force coefficient of the energy dissipation system at the first (ground) story. Using Equation (14) and noting that η_i = η, then ${}_s\overline{\alpha }{}_i$ is given by:

$${}_s\overline{\alpha }{}_i=\frac{M}{{{\displaystyle \sum }}_{j=i}^N{m}_j}\sqrt{\frac{K_i{}_{f}k{}_i{{\displaystyle \sum }}_{n=1}^r\left({\psi }_{mn,i}{\mathsf{\Gamma }}_n^2{E}_{hn}^{SDOF}\right)}{K_1{}_{f}k{}_1{{\displaystyle \sum }}_{n=1}^r\left({\psi }_{mn,1}{\mathsf{\Gamma }}_n^2{E}_{hn}^{SDOF}\right)}}$$

(15)

Taking into account that _sδ_y,i = _sQ_y,i/_sk_i, recalling Figure 1 and noting that Q_y,i = _sQ_y,i+ _fk_{i s}δ_y,i =_sQ_y,i+ (_sk_i/K_i) _sδ_y,i = _sQ_y,i+_sQ_y,i(1/K_i) = _sQ_y,i(K_i + 1)/K_i, the yield shear force coefficient of the whole structure (including main structure and energy dissipation system) at the i-th story α_i is related to _sα_i by:

$${\alpha }_i={}_s\alpha {}_i\left(\frac{K_i+1}{K_i}\right)$$

(16)

Therefore, the distribution of the optimum total yield shear force coefficient, ${\overline{\alpha }}_i={\alpha }_{opt,i}/{\alpha }_1$, is simply ${\overline{\alpha }}_i={\alpha }_{opt,i}/{\alpha }_1=\left[{}_s\alpha {}_{opt,i}\left(K_i+1\right)/K_i\right]/\left[{}_s\alpha {}_1\left(K_1+1\right)/K_1\right]={}_s\overline{\alpha }{}_i\left[K_1\left(K_i+1\right)\right]/\left[K_i\left(K_1+1\right)\right]$. If K_i is the same in all stories, it is obvious that ${\overline{\alpha }}_i$ coincides with ${}_s\overline{\alpha }{}_i$. Using Equation (15) ${\overline{\alpha }}_i$ gives, then:

$${\overline{\alpha }}_i=\frac{M}{{{\displaystyle \sum }}_{j=i}^N{m}_j}\sqrt{\frac{K_i{}_{f}k{}_i{{\displaystyle \sum }}_{n=1}^r\left({\psi }_{mn,i}{\mathsf{\Gamma }}_n^2{E}_{hn}^{SDOF}\right)}{K_1{}_{f}k{}_1{{\displaystyle \sum }}_{n=1}^r\left({\psi }_{mn,1}{\mathsf{\Gamma }}_n^2{E}_{hn}^{SDOF}\right)}}\frac{K_1\left(K_i+1\right)}{K_i\left(K_1+1\right)}$$

(17)

Past research [7] shows that in an elastic-perfectly plastic MDOF system designed with the optimum distribution associated with a given ground motion, damage spreads out evenly among the stories essentially regardless of the level of plastic deformation. This allows to assume that if the ground motion is scaled down by a factor s_y so that the level of plastic deformation in the dampers is zero, i.e., η_i = η ≈ 0, the maximum total elastic force developed by the dampers at each story i in each mode n, _sQ_maxn,i, coincides with the lateral strength required on the dampers in each mode n, _sQ_yn,i, to carry all stories to the brink of yielding, that is _sQ_maxn,i = _sQ_yn,i. These forces can be expressed in terms of shear force coefficients, _sα_max,i, and their distribution, i.e., ${}_s\overline{\alpha }{}_{max,i}={}_s\alpha {}_{max,i}/{}_s\alpha {}_{max,1}$, would coincide with the optimum distribution ${}_s\overline{\alpha }{}_i$, i.e., ${}_s\overline{\alpha }{}_{max,i}={}_s\overline{\alpha }{}_i$. The vector of maximum elastic forces on the whole structure (i.e., main structure and energy dissipation system) at the n-th vibration mode, F_n, of an elastic MDOF system subjected to a given ground motion characterized by its elastic response spectrum S_a-T scaled by factor s_y can be easily obtained from modal analysis as follows:

$${\mathbf{F}}_n=\mathbf{M}{\phi }_{\mathit{n}}{\mathsf{\Gamma }}_n{s}_y{S}_{an}$$

(18)

where S_an is the ordinate of the S_a-T spectrum at the period of vibration mode n. The corresponding maximum total shear force on the i-th story in mode n, Q_maxn,i, is simply obtained by summing up the components F_n,_i of vector F_n above the i-th story as follows:

$$Q_{maxn,i}=s_y{\mathsf{\Gamma }}_{\mathrm{n}}{S}_{an}{\displaystyle \sum }_{j=i}^N\left(m_j{\phi }_{n,j}\right)$$

(19)

where φ_n,j are the components of mode vector φ_n. The maximum force in the energy dissipation system at the i-th story in mode n, _sQ_maxn,i, can be related with Q_maxn,i by $Q_{maxn,i}={}_{s}Q{}_{maxn,i}+\left({}_{s}Q{}_{maxn,i}/{}_{s}k{}_i\right){}_{f}k{}_i={}_{s}Q{}_{maxn,i}\left(K_i+1\right)/K_i$. Solving for _sQ_maxn,i in this expression and recalling that is has been assumed that _sQ_maxn,i = _sQ_yn,i, the following expression is obtained:

$${}_{s}Q{}_{yn,i}=\frac{K_i}{\left(1+K_i\right)}{s}_y{\mathsf{\Gamma }}_{\mathrm{n}}{S}_{an}{\displaystyle \sum }_{j=i}^N\left(m_j{\phi }_{n,j}\right)$$

(20)

In turn, for a design earthquake that causes plastic deformations in the energy dissipation system (i.e., a ground motion not scaled down by s_y), the maximum and minimum normalized plastic deformations experienced by the energy dissipation system of each story i in each vibration mode n are defined as follows: μ_maxn,i = (δ_maxn,i − _sδ_yn,i)/_sδ_yn,i and μ_minn,i = |(δ_minn,i − _sδ_yn,i)|/_sδ_yn,i. Here δ_maxn,i and δ_minn,i are the maximum and minimum inter-story drifts attained at the i-th story in mode n, and _sδ_yn,i is defined as _sδ_yn,i = _sQ_yn,i/_sk_i. The mean value is ${\overline{\mu }}_{n,i}=\left({\mu }_{maxn,i}+{\mu }_{minn,i}\right)/2$. It has been shown [7] that, in mixed systems combining in each story an elastic part with an elastic-perfectly plastic part, as shown in Figure 1, the maximum deformations are roughly the same in the positive and negative domains and thus it can be assumed δ_maxn,i ≈ |δ_minn,i|, or ${\mu }_{maxn,i}\approx {\mu }_{minn,i}$. Moreover, it has been shown [7] that if the structure is designed with the optimum distribution, then ${\overline{\mu }}_{n,i}$ is roughly the same in all stories and thus it can be assumed that ${\overline{\mu }}_{n,i}={}_s\mu {}_n$. Consequently (δ_maxn,i − _sδ_y,i) = _sμ_{n s}δ_y,i. The hysteretic energy accumulated at the i-th story, if pushed monotonically beyond the elastic range following the distribution of lateral forces of the n-th mode given by Equation (18) up to δ_maxn,i, is:

$$E_{hmn,i}={}_{s}Q{}_{yn,i}\left({\delta }_{maxn,i}-{}_s\delta {}_{yn,i}\right)={}_{s}Q{}_{yn,i}{}_s\mu {}_n{}_s\delta {}_{yn,i}$$

(21)

Recalling that _sδ_yn,i = _sQ_yn,i/_sk_i and K_i =_sk_i/_fk_i, using Equations (12), (20), and (21) ψ_mn,i can estimated with the following expression:

$${\psi }_{mn,i}=\frac{\frac{K_i}{{}_{f}k{}_i{\left(1+K_i\right)}^2}{\left({{\displaystyle \sum }}_{j=i}^N{m}_j{\phi }_{n,j}\right)}^2}{{{\displaystyle \sum }}_{k=1}^N\left[\frac{K_k}{{}_{f}k{}_k{\left(1+K_k\right)}^2}{\left({{\displaystyle \sum }}_{j=k}^N{m}_j{\phi }_{n,j}\right)}^2\right]}$$

(22)

Equation (22) indicates that when the distribution of lateral strength of the structure, expressed in terms of the yield-shear force coefficient, follows the optimum distribution, then ψ_mn,i can be calculated from the mass and the elastic properties of the structures, with no need to perform a pushover analysis. In this study, Equation (15) with ψ_mn,i calculated using Equation (22) is proposed as the optimum yield-shear force coefficient distribution for the energy dissipation system and will be denoted as ${}_s\overline{\alpha }{}_{prop,i}$ herein. Similarly, Equation (17) with ψ_mn,i calculated using Equation (22) is proposed as the optimum yield-shear force coefficient distribution for the whole structure and will be denoted as ${\overline{\alpha }}_{prop,i}$ hereafter.

4. Numerical Validation

4.1. Prototype Structures, Modelization and Ground Motions

In order to validate the optimum yield shear-force coefficient distribution ${\overline{\alpha }}_{prop,i}$ proposed in this study, several prototype reinforced concrete (RC) structures consisting of waffle-flat plates supported on RC columns having 3, 6, and 9-stories (main structure) equipped with metallic dampers (energy dissipation system) were investigated. To model the prototype structures, it was assumed that the mass of each plate is concentrated in one point of mass m_i located at the centre of mass of the plate (lumped-mass model). Torsional and rocking effects are not considered and therefore a single degree of freedom is assigned to each mass: the horizontal displacement in the direction of the ground motion. The rest of parameters that define the numerical model are the i-th story lateral stiffness of the main structure _fk_i; and the lateral stiffness _sk_i and lateral strength _sQ_yi provided by the dampers. _fk_i is defined as the quotient between the shear force endured by the main structure at the i-th story and the corresponding interstory drift in elastic conditions. _sk_i is defined as the quotient between the shear force endured by the metallic dampers installed at the i-th story and the corresponding interstory drift in elastic conditions. _sQ_yi is the maximum shear force that can be endured by the metallic dampers installed at the i-th story. The main structure is assumed to remain elastic and the metallic dampers exhibit an elastic-perfectly plastic hysteretic behaviour, as shown in Figure 1. The inherent damping fraction ξ = 0.05. The main structures are designed to sustain only the gravity loads specified under Spanish standards [21,22]. Therefore, _fk_i and m_i for the prototypes with the same number of stories are identical. _fk_i was determined from a static analysis in which the main structure was subjected to a distribution of lateral forces that followed the pattern of the first vibration mode.

Table 1. Lateral stiffness, mass and height of the main structure.

Prototype	3-Story			6-Story			9-Story
Story	fki (kN/cm)	mi (kNs2/cm)	hi (m)	fki(kN/cm)	mi (kNs2/cm)	hi (m)	fki(kN/cm)	mi (kNs2/cm)	hi (m)
9	-	-	-	-	-	-	1523	5.24	3.10
8	-	-	-	-	-	-	1572	5.68	3.10
7	-	-	-	-	-	-	1618	5.68	3.10
6	-	-	-	1071	3.51	3.10	2149	5.68	3.10
5	-	-	-	1120	3.89	3.10	2191	5.68	3.10
4	-	-	-	1126	3.89	3.10	2206	5.68	3.10
3	571	2.22	3.10	1141	3.89	3.10	2707	5.68	3.10
2	569	2.56	3.10	1518	3.89	3.10	2803	5.68	3.10
1	510	2.56	3.50	1636	3.89	3.50	3259	5.68	3.50
fT1 (s)	0.94			1.38			1.81

shows _fk_i, m_i, the height of the story i h_i and the fundamental period of the main structure _fT₁. To validate the proposed optimum distribution in a large number of different cases, a variety of levels of plastic deformation for the dampers—expressed in terms of coefficient _sμ—, and a variety of ground motions—characterized by the seismological parameters V_D, T_G, T_NH, I_d and the proximity to the source (i.e., near-fault or far-field earthquakes)—were considered. T_G is the predominant period of the ground motion, T_NH is the corner period of the pseudo-velocity Newmark-Hall spectrum, and I_D is a seismological parameter proposed by Cosenza and Manfredi [23] defined by I_D = 2gI_A/(π PGA PGV) where I_A is the Arias intensity, PGA is the peak ground acceleration and PGV the peak ground velocity. Appendix A presents the original (unscaled) ground motions selected, together with T_G, T_NH, I_D, PGA and PGV. The lateral stiffness provided by the dampers at each story i, _sk_i, was made proportional to the lateral stiffness of the main frame _fk_i, i.e., K_i = _fk_i/_sk_i= K. Each ground motion was scaled in acceleration by applying a factor SF. The values of SF were kept in the range 0.3 ≤ SF ≤ 3 to distort as little as possible the characteristics of the original signal, and were determined to keep the maximum allowed inter-story drift IDI_,i below 0.75%, while providing a K smaller than 20. The limit IDI_,i ≤ 0.75% was fixed on the basis of code recommendations [8,24] and experimental results obtained by the authors in previous studies [25,26,27]. Also, values of K larger than 20 are not realistic because they lead to very small values of _sδ_yi that can hardly be applied in practice. For a given set of values of parameters _sμ, SF, V_D, T_G, T_NH, I_d and the proximity to the source, the required K, _sQ_y1—expressed in terms of the base shear force coefficient _sα₁ (=_sQ_y1/Mg)—, and the normalized energy dissipation demand of the dampers η were calculated applying the method developed by Benavent–Climent [12], but using the optimum yield shear force distribution proposed in this study, with Equations (15) and (22) instead of the optimum distribution proposed in [12].

Table 2. Properties of the designed metallic in three-story buildings.

Records	SF	K	T1 (s)	VD (cm/s)	sα1	η	sμ
Near-Field	-	-	-	-	-	-	-
El Centro	0.90	6.6	0.34	65	0.18	15.7	5.7
Kobe	0.35	8.1	0.31	66	0.18	19.0	7.1
Lorca	1.00	8.5	0.31	53	0.22	7.5	5.9
Tolmezzo	1.25	8.3	0.31	60	0.18	16.1	7.4
Korinthos	1.00	4.1	0.42	60	0.16	10.3	3.8
Duzce (Duzce)	0.75	11.0	0.27	79	0.18	42.3	10.4
Kalamata	1.00	4.7	0.39	49	0.16	7.0	4.6
Far-Field	-	-	-	-	-		-
Duzce (Izmit)	1.00	9.6	0.29	55	0.18	14.4	8.6
Montebello	1.75	12.2	0.26	65	0.17	30.8	12.0
Petrovac	0.50	7.7	0.32	80	0.16	37.3	7.8
Hachinoe	0.95	7.7	0.32	53	0.18	10.7	6.9
Taft	1.00	1.8	0.56	52	0.07	16.3	3.8
Calitri	1.00	1.6	0.58	60	0.05	38.1	4.8
Tabas	0.40	6.0	0.36	55	0.18	9.0	5.0

,

Table 3. Properties of the designed metallic in six-story buildings.

Records	SF	K	T1 (s)	VD (cm/s)	sα1	η	sμ
Near-Field	-	-	-	-	-	-	-
El Centro	0.70	3.7	0.64	65	0.13	6.6	2.9
Kobe	0.30	5.9	0.52	69	0.14	11.3	4.9
Lorca	1.00	6.0	0.52	46	0.06	18.1	12.9
Tolmezzo	1.58	4.6	0.58	66	0.17	5.0	2.9
Korinthos	1.20	4.5	0.59	69	0.15	6.9	3.1
Duzce (Duzce)	0.75	7.9	0.46	83	0.14	23.3	6.9
Kalamata	1.30	17.7	0.32	65	0.17	18.9	13.3
Far-Field	-	-	-	-	-	-	-
Duzce (Izmit)	1.00	12.2	0.38	65	0.17	12.9	8.8
Montebello	2.30	7.5	0.47	72	0.18	9.4	5.0
Petrovac	0.62	2.0	0.80	64	0.07	13.5	3.3
Hachinoe	1.10	3.4	0.66	53	0.14	3.2	2.4
Taft	1.70	8.2	0.45	81	0.14	22.6	7.3
Calitri	1.00	2.9	0.70	76	0.06	32.3	5.3
Tabas	0.40	6.8	0.49	63	0.12	12.8	7.1

and

Table 4. Properties of the designed metallic in nine-story buildings.

Records	SF	K	T1 (s)	VD (cm/s)	sα1	η	sμ
Near-Field	-	-	-	-	-	-	-
El Centro	0.90	7.8	0.61	85	0.15	10.9	5.2
Kobe	0.36	10.7	0.53	86	0.15	15.7	7.6
Lorca	1.60	6.0	0.68	62	0.10	9.3	7.3
Tolmezzo	1.80	5.0	0.74	76	0.15	5.3	3.2
Korinthos	1.50	16.7	0.43	90	0.15	28.1	12.5
Duzce (Duzce)	0.90	2.9	0.92	79	0.10	7.8	2.8
Kalamata	1.20	6.0	0.68	53	0.05	25.4	15.7
Far-Field	-	-	-	-	-	-	-
Duzce (Izmit)	1.10	17.6	0.42	75	0.15	18.9	13.2
Montebello	3.30	5.0	0.74	74	0.12	8.2	4.4
Petrovac	0.80	4.2	0.79	87	0.10	14.3	4.3
Hachinoe	1.20	7.1	0.64	67	0.12	8.9	6.3
Taft	1.85	2.1	1.03	69	0.06	9.6	3.2
Calitri	1.00	4.1	0.80	92	0.08	20.8	4.7
Tabas	0.35	8.8	0.58	75	0.15	9.5	6.3

summarize the combinations of values for SF, K, V_D, _sα₁, η, _sμ investigated, together with the resulting fundamental period of the whole structure in the elastic range T₁. In applying Equation (15), the input energy $E_{hn}^{SDOF}$ corresponding to each mode n was determined from the inelastic hysteretic energy spectrum E_h-T of the scaled ground motion.

The inelastic hysteretic energy spectrum E_h-T was calculated by following Akiyama’s approach [7], as explained next. First, the elastic input energy spectrum E_I-T of the ground motion was obtained with the expression of the right-hand side of Equation (3). Second, the inelastic input energy spectrum, ${\overline{E}}_I-T$, was determined by averaging the elastic spectrum E_I-T in the range T ≤ T ≤ T_max; that is, the ${\overline{E}}_I$ corresponding to a given period T of the spectrum was obtained by ${\overline{E}}_I={{\displaystyle \int }}_T^{T_{max}}{E}_{I}dT/\left(T_{max}-T\right)$. Here, T_max is a period longer than T, influenced by the level of plastic deformations experienced by the system characterized by _sμ. The calculation of T_max is explained in Appendix B as an extension of the formulation proposed by Akiyama [7] for conventional structures to systems with metallic dampers. Third, recalling that the elastic vibrational energy E_e can be neglected as explained in Section 2.2. (i.e., E_h = E_D), E_h is readily obtained from ${\overline{E}}_I$ by means of Equation (6) as $E_h={\overline{E}}_I/{\left(1+3\mathsf{\xi }+1.2\surd \mathsf{\xi }\right)}^2$.

4.2. Optimum Distributions ${\overline{\alpha }}_{prop,i}$ Obtained with the Proposed Expression for Different Ground Motions

Figure 2 and Figure 3 show the proposed optimum yield shear force coefficient distributions ${\overline{\alpha }}_{prop,i}$—Equations (17) and (22)—obtained for the near-fault (Figure 2) and far-field (Figure 3) ground motions. In all cases the grow is approximately linear in the lower two thirds of the building, and increases exponentially in the upper stories, especially for taller buildings. This is due to the influence of the higher modes of vibration. It is also worth noting that in the lower two thirds part of the building ${\overline{\alpha }}_{prop,i}$ is basically the same for all records, whereas clear differences are observed in the upper one third depending on the ground motion considered. These tendencies and features have been identified in previous studies [7,12].

4.3. Comparison between Proposed and “Exact” Distributions

Similarly to the optimum yield shear force coefficient distributions proposed in previous studies [7,11,12] the new distribution proposed in this study provides an approximation to the “exact” distribution, referred to as ${\overline{\alpha }}_{exact,i}$ hereafter, that would make η_i exactly the same in all stories. ${\overline{\alpha }}_{exact,i}$ is different for each ground motion and can be obtained only through an iterative trial-and-error procedure of nonlinear time history analyses. A number of multi-purpose algorithms existing in the literature may be used to efficiently conduct such an iterative process [28,29]. In this study, ${\overline{\alpha }}_{exact,i}$ was calculated using the pattern search method (PSM), which belongs to the direct search methodology for optimization of n-dimensional functions [30]. PSM has been used in previous studies on optimization of buildings with dampers [31] and on damage control of buildings [32,33]. There are other approaches such as the gradient-based method or genetic algorithms used successfully in structural optimization problems [28,29] that could potentially be more efficient to reach the “exact” optimum distribution. Nevertheless, it is out of the scope of this study to determine the most efficient method. In short, PSM is an iterative procedure to minimize an objective function. Some advantages of the PSM are: (i) it is a derivative-free method and (ii) it is simple. In this case, the objective function is the standard deviation of η_i, while the independent n-dimensional variable is the lateral strength of the structure. Through non-linear time history analyses with a given accelerogram, PSM varies the lateral strength distribution of the structure until the standard deviation of η_i is zero. The main drawback of using PSM in this study has been the significant increase of the calculation time with the increase of the number of stories of the buildings. PSM took about fifteen minutes to reach the optimum distribution for a three-story building subjected to a single ground motion. This time increased up to four hours for a six-story building and up to ten hours for a nine-story building.

Using the PSM, the exact distribution ${\overline{\alpha }}_{exact,i}$ was calculated for each prototype structure and for each ground motion, and then was compared with the counterpart approximate distribution proposed in this study, ${\overline{\alpha }}_{prop,i}$ in Figure 4 and Figure 5. The figures show in bold lines the mean curves obtained averaging ${\overline{\alpha }}_{prop,i}/{\overline{\alpha }}_{exact,i}$, by means of the seven accelerograms applied to each prototype structure at each story level. It can be seen that the mean curves of ${\overline{\alpha }}_{prop,i}/{\overline{\alpha }}_{exact,i}$ are very close to the vertical line passing through the abscissa 1.0 (ranging from 0.9 to 1.2); this indicates that the proposed distribution is very close to the exact distribution. For comparison purposes, Figure 4 and Figure 5 also show the mean curves obtained using the optimum distribution proposed by Benavent-Climent [12], ${\overline{\alpha }}_{Benavent,i}$ normalized by ${\overline{\alpha }}_{exact,i}$. It is clear in the figures that while both distributions ${\overline{\alpha }}_{prop,i}$ and ${\overline{\alpha }}_{Benavent,i}$ approximate the exact solution in the lower part of the structure, in the upper part the distribution proposed by Benavent-Climent deviates up to about 35% from the exact value.

Under near-field ground motions, the coefficients of variation (COV) of ${\overline{\alpha }}_{prop,i}/{\overline{\alpha }}_{exact,i}$ calculated for all stories are 0.05, 0.07 and 0.10 for the three-, six-, and nine-story prototype structures, respectively. The counterpart COVs for far-field ground motions are 0.05, 0.06 and 0.08, respectively. Noticeably, there are not significant differences between the two types of earthquakes, the COV of ${\overline{\alpha }}_{prop,i}/{\overline{\alpha }}_{exact,i}$ being less than or equal than 0.10, independently the height of the building. The optimum distribution proposed by Benavent-Climent [12] provides a good approximation to the “exact” distribution in the lower part of the building but tends to over-estimate the required strength in the upper stories. The COV of ${\overline{\alpha }}_{Benavent,i}/{\overline{\alpha }}_{exact,i}$ under near-field records are 0.14, 0.14, 0.17 for three, six and nine-story prototype structures respectively, and 0.13, 0.13, 0.17 under far-field records. Therefore, the proposed optimum distribution ${\overline{\alpha }}_{prop,i}$ gives a better approximation to the “exact” one than to that obtained with ${\overline{\alpha }}_{Benavent,i}$. It is worth noting that, from a practical point of view, the main differences between ${\overline{\alpha }}_{Benavent,i}$ and ${\overline{\alpha }}_{prop,i}$, are: (i) ${\overline{\alpha }}_{Benavent,i}$ is based only on properties of the main structure while ${\overline{\alpha }}_{prop,i}$ requires also to know the ratio between the lateral stiffness of the dampers and the lateral stiffness of the main frame at each story; (ii) ${\overline{\alpha }}_{Benavent,i}$ depends on the predominant period of the ground motion T_G while ${\overline{\alpha }}_{prop,i}$ requires to know the inelastic input energy spectrum. Therefore, the time required to obtain the optimum distribution with ${\overline{\alpha }}_{prop,i}$ is larger than with ${\overline{\alpha }}_{Benavent,i},$ and the analytical expression of the former is more complex than the later. However, from a practical point of view, since these calculations are made usually with a computer, the differences in time and in programming complexity are negligible.

4.4. Damage Distribution among Stories

As explained in previous sections, the damage in a given story i can be characterized by the ratio η_i. It has likewise been explained in previous sections that the optimum distribution of the shear force coefficients is the one that makes η_i equal in all stories. This strictly occurs only when the distribution ${\overline{\alpha }}_{exact,i}$ is used, and the common value of η_i in this case will be named η_exact (=η_i) hereafter. When the approximate distributions ${\overline{\alpha }}_{prop,i}$ or ${\overline{\alpha }}_{Benavent,i}$ are used, in general the η_i’s do not have exactly the same value in all stories. In order to evaluate such deviations, the η_i obtained at each story through nonlinear time story analyses conducted for each individual ground motion and for each prototype structure designed with ${\overline{\alpha }}_{prop,i}$, referred to as η_prop,i hereafter, are plotted in Figure 6 and Figure 7 normalized by η_exact. Also shown in the figures with bold lines are the mean curves of η_prop,i/η_exact obtained by averaging the response at each story level throughout the seven ground motions applied to each prototype structure. As can be seen, the mean curve η_prop,i/η_exact is close to the vertical line passing through the abscissa 1.0. This indicates that the proposed distribution provides a reasonably uniform distribution of damage among stories. More precisely, the COV of η_prop,i/η_exact calculated for all stories and all the near-field earthquakes are 0.40, 0.33 and 0.42, for the three-, six-, and nine-story prototype structures, respectively; the counterpart values for far-field earthquakes are 0.33, 0.33, and 0.37.

For comparison, the η_i obtained at each story through nonlinear time story analyses conducted for each individual ground motion and for each prototype structure designed with ${\overline{\alpha }}_{Benavent,i}$ was also calculated and it is referred to as η_Benavent,i hereafter. Figure 6 and Figure 7 show in bold lines the mean curves of η_Benavent,i/η_exact obtained by averaging the response at each story level through the seven ground motions applied to each prototype structure. When ${\overline{\alpha }}_{Benavent,i}$ is used the COV of η_Benavent,i/η_exact are 0.81, 0.61, and 0.63 for near-field earthquakes and 0.80, 0.59, and 0.52 for far-field earthquakes, respectively. It follows from these results that the optimum yield shear force distribution proposed in this study provides a much more uniform distribution of damage among the stories than the one proposed by Benavent-Climent [12]. The reason lies in the bias of Benavent-Climent’s equation, overestimating the required strength at the higher stories, which leads to lower values of η_Benavent,i for the high stories, and higher values for the low stories.

Furthermore, Figure 8 compares the histories of energy E_h dissipated by the 6-story building designed with ${\overline{\alpha }}_{prop,i}$ and with ${\overline{\alpha }}_{Benavent,i}$ for two ground motions: Calitri (Figure 8a) and Montebello (Figure 8b). It can be seen that there are not significant differences between them.

4.5. Maximum Interstory Drifts

Figure 9 and Figure 10 show the maximum inter-story drifts normalized by the story height, IDI_i, obtained from the time history analyses conducted on the prototype structures designed with ${\overline{\alpha }}_{prop,i}$, denoted by IDI_prop,i (record-to-record, mean and the envelope of the maximum values “maxenv”), and for those designed with ${\overline{\alpha }}_{exact,i}$, denoted by IDI_exact,i (the mean and the envelope of the maximum values “maxenv”). Overall, comparing the mean curves it is seen that the IDI_i’s obtained with the proposed distribution (i.e., “mean IDI_prop,I” in the figures) are very close, almost coinciding with the exact values (i.e., “mean IDI_exact,i” in the figures). Similar observations can be made with the maximum envelope curves (i.e., “maxenv IDI_prop,I” versus “maxenv IDI_exact,I”). Moreover, the target IDI_i (=0.75%) is not exceeded in any case. It is also noteworthy that despite the differences observed in the distribution of η_i for the structures designed with ${\overline{\alpha }}_{prop,i}$ and with ${\overline{\alpha }}_{exact,i}$ (Figure 6 and Figure 7), the response in terms of displacement are in very good agreement. In other words, the small deviations of η_i among stories scarcely affect the maximum lateral inter-story drifts.

Deviations of the yield shear force coefficient distribution with respect to the exact value ${\overline{\alpha }}_{exact,i}$ inevitably cause deviations from the aspired even distribution of damage (i.e., η_i equal in all stories). Nonetheless, when these deviations are limited (to about 20% as occurs with the proposed distribution ${\overline{\alpha }}_{prop,i}$ according to the results of this study), then the damage concentration (i.e., values of η_i in specific stories markedly larger than in the rest of stories) is prevented and the response of the structure in terms of displacements IDI_prop,i scarcely deviates from that obtained for a structure designed ${\overline{\alpha }}_{exact,i}$, i.e from IDI_exact,i. These results emphasize the key importance of designing structures with a strength distribution close to that prescribed by the optimum yield-shear force distribution so as to prevent damage concentration and to control the response of the structure in terms of displacement. The results of this study also underline the improvements of the proposed optimum distribution with respect to the one proposed by Benavent-Climent [12] in past studies.

Finally, a comparison was carried out between the IDI_i’s of the prototype structures designed using ${\overline{\alpha }}_{prop,i}$ or with ${\overline{\alpha }}_{Benavent,i}$, respectively IDI_prop,i and IDI_Benavent,i. The goodness of the agreement of IDI_i or IDI_Benavent,i with respect to IDI_exact,i, is quantified in terms of the normalized square mean error, NMSE, defined as follows:

$$NMSE=1-\frac{{{\displaystyle \sum }}_{i=1}^N{\left(ID{I}_i-ID{I}_{exact,i}\right)}^2}{{{\displaystyle \sum }}_{i=1}^N{\left(ID{I}_{exact,i}-\mathrm{mean}(ID{I}_{exact,i})\right)}^2}$$

(23)

NMSE ranges between −∞ (very bad fit) and 1 (perfect fit). Values above 0.70 are considered a reasonably good fit. The results are shown in Figure 11. It is seen that the structures designed using ${\overline{\alpha }}_{prop,i}$ lead to NMSE values greater than 0.70 (Figure 11a) and 0.50 (Figure 11b) in the case of near-field and far-field records, respectively. Nonetheless, for the structures designed with ${\overline{\alpha }}_{Benavent,i}$, even though the target IDI_i is not exceeded and most analyses show NSME greater than 0.50, there are results giving poor values for NMSE, especially with near-field records (Figure 11a).

4.6. Ductility Distribution among Stories

Following past studies [7], it was assumed in Section 3 that when the structure is designed with the optimum yield-force coefficient distribution the ductility demand ${\overline{\mu }}_i$ is rougly the same in all stories (i.e., ${\overline{\mu }}_i={}_s\mu$). This assumption is further examined in this subsection. Using the IDI_prop,i’s of Section 4.5, the corresponding ductility factor ${\overline{\mu }}_i$ at each story i, referred to as ${\overline{\mu }}_{prop,i}$ herein, was calculated for the structures designed with the proposed optimum distribution ${\overline{\alpha }}_{prop,i}$ by:

$${\overline{\mu }}_{prop,i}=[(h_{i}ID{I}_{prop,i}/100)-{}_s\delta {}_{yi}]/{}_s\delta {}_{yi}$$

(24)

where h_i is given in Table 1. The yield displacement of the dampers is obtained with:

$${}_s\delta {}_{yi}={}_{s}Q{}_{yi}/{}_{s}k{}_i=[{\overline{\alpha }}_{prop,i}{}_s\alpha {}_1{{\displaystyle \sum }}_{j=i}^N{m}_{j}g]/[K_i{}_{f}k{}_i]$$

(25)

where m_i, ${}_{f}k{}_i$ are given in Table 1 and ${}_s\alpha {}_1$, K_i = K in Table 2, Table 3 and Table 4. ${\overline{\mu }}_{prop,i}$ was compared with the ductility factor ${}_s\mu$ assumed for designing the prototype structures shown in Table 2, Table 3 and Table 4. The quotient ${\overline{\mu }}_{prop,i}/{}_s\mu$ calculated for each ground motion and the mean curve “mean ${\overline{\mu }}_{prop,i}/{}_s\mu ”$ are plotted in Figure 12 and Figure 13. Also plotted in the figures are the mean curves “mean ${\overline{\mu }}_{Benavent,i}/{}_s\mu$” and “mean ${\overline{\mu }}_{exact,i}/{}_s\mu$” obtained using IDI_Benavent,i, ${\overline{\alpha }}_{Benavent,i}$ and IDI_exact,i, ${\overline{\alpha }}_{exact,i}$ instead of IDI_prop,I and ${\overline{\alpha }}_{prop,i}$ in Equations (24) and (25). The curves “mean ${\overline{\mu }}_{prop,i}/{}_s\mu$”, “mean ${\overline{\mu }}_{Benavent,i}/{}_s\mu$” and “mean ${\overline{\mu }}_{exact,i}/{}_s\mu$” are found to lie approximately in a vertical line. Noting that the denominator ${}_s\mu$ is the same in all cases and has the same value for all stories, this means that the ductility demand ${\overline{\mu }}_i$ is approximately the same in all stories. The variation of “mean ${\overline{\mu }}_{prop,i}/{}_s\mu ”$, “mean ${\overline{\mu }}_{Benavent,i}/{}_s\mu$” or “mean ${\overline{\mu }}_{exact,i}/{}_s\mu$” throughout the stories was further quantified by calculating the corresponding COV. The results are shown in

Table 5. COV of “mean ${\overline{\mu }}_{exact,i}/{}_s\mu$”, “mean ${\overline{\mu }}_{prop,i}/{}_s\mu$” and “mean ${\overline{\mu }}_{Benavent,i}/{}_s\mu$”.

Type of Records	Near-Field			Far-Field
Prototypes	N			N
	3	6	9	3	6	9
COV of “mean ${\overline{\mu }}_{exact,i}/{}_s\mu$”	0.10	0.08	0.12	0.05	0.07	0.11
COV of “mean ${\overline{\mu }}_{prop,i}/{}_s\mu$”	0.35	0.18	0.16	0.18	0.14	0.16
COV of “mean ${\overline{\mu }}_{Benavent,i}/{}_s\mu$”	0.69	0.43	0.32	0.36	0.24	0.28

. It can be seen that the structures designed with ${\overline{\alpha }}_{exact,i}$ so that η is exactly the same in all stories also exhibit a very uniform distribution of ${\overline{\mu }}_i$ among stories, the COVs ranging from 0.05 to 0.12. This confirms Akiyama’s [7] results and further supports one of the assumptions made in this study. For structures designed with the proposed approximate optimum distribution ${\overline{\alpha }}_{prop,i}$, the COV slightly increases with respect to ${\overline{\alpha }}_{exact,i}$, in particular for the three story structure, but is still small. The distribution proposed by Benavent-Climent [12] presents the largest COVs.

Finally, it is likewise observed in Figure 12 and Figure 13 that “mean ${\overline{\mu }}_{prop,i}/{}_s\mu ”$, “mean ${\overline{\mu }}_{Benavent,i}/{}_s\mu$” and “mean ${\overline{\mu }}_{exact,i}/{}_s\mu$” are in general below 1. The reason is that the expression used in reference [12] to relate ${\overline{\mu }}_i$ with ${\eta }_i$, developed by Cosenza and Manfredi [23], is somewhat conservative and provides in general a safe-side estimation of the maximum displacements.

5. Experimental Validation

5.1. Test Model

A waffle-flat slab RC structure supported on RC columns and equipped with metallic dampers (test model) was built in the Laboratory of Structures of the University of Granada (Spain), and subjected to unidirectional dynamic loadings with a shaking table. The test model is representative of a prototype two story building at a scale of two fifths. The overall dimensions and characteristics of the test model are shown in Figure 14. Due to space limitations in the laboratory, the test model does not represent the whole buildings but rather a portion of it. More precisely, to reduce its height it was assumed that under lateral loadings the sections of zero bending moment on the columns of the second story are located at mid-height. With this assumption, the columns of the second story were replaced by half columns with a pin connection at the top. The weight that the prototype building supports above mid-height of these columns was replaced by steel blocks put on the top (indicated as “added weight” in Figure 14). Similarly, to reduce the horizontal dimensions of the test model it was assumed that under lateral loading the mid-span section of the RC slab moves horizontally (i.e., there is no vertical movement) and that the bending moment is zero. To simulate these boundary conditions, the mid-span section of the RC slab was connected to the steel blocks put on the top of the half-columns of the second story with pin-ended steel bars. The very large flexural stiffness of these steel blocks prevented any possible vertical displacement. Steel blocks were also attached to the slab to represent the gravity loads acting on the floor. The mass m₁ of the RC floor slab with the added weight was m₁ = 4079 kg. The mass m₂ of the added weight put on top of the half columns of the second story was m₂ = 7058 kg. The total mass of the test model was 11137 kg.

One metallic damper was installed in each story along the direction of shaking. The type of metallic damper used was developed by the authors in past studies [34] and is named web plastification damper (WPD). The details of the WPDs are shown in Figure 15. The WPD is constructed by assembling several short-length segments of I-shaped steel sections and two U-shaped steel bars. The WPD is installed in the structure as a diagonal bar connecting two consecutive floor levels. When the damper is subjected to axial deformations, the webs of the I-sections are forced to undergo out of plane plastic deformations. The two U-shaped steel bars function as auxiliary elements aimed to transfer the horizontal forces from the slabs to the I-shaped steel sections and are designed to remain elastic.

5.2. Test Set-Up and Instrumentation

The RC elements of the test model were instrumented with strain gages attached to the longitudinal reinforcement, uniaxial accelerometers and displacement transducers (linear variable differential transformers, LVDT). The ends of the WPDs were also instrumented with strain gages that measured the axial forces acting on them. The accelerometers and the LVDT’s provided the horizontal accelerations and displacements of the masses. Additional LVDTs installed in the WPDs measured their axial displacements. Data were acquired continuously with a scan frequency of 200 Hz. Figure 16 offers a general view of the experimental set up.

5.3. Seismic Simulations

The test model was subjected to three consecutive seismic simulations. In each simulation, the shaking table reproduced the accelerogram recorded at Calitri during the Campano-Lucano earthquake (Italy, 1980) scaled in time by the factor 0.63. In the first seismic simulation, C100, the original accelerogram was applied without scaling in amplitude. In the second simulation, C200, the amplitude of the original accelerogram was multiplied by two, and in the third simulation C300 by three. The resulting peak accelerations applied to the shaking table were 0.16, 0.31, and 0.47 g, respectively (here g is the acceleration of gravity). Figure 17a shows the accelerogram (scaled only in time by 0.63) and Figure 17b the corresponding elastic spectrum E_I-T corresponding to a damping fraction of ξ = 1.8% (i.e., the damping fraction measured during the test as discussed later).

5.4. Test Results

The readings provided by the gauges attached to the steel reinforcement of the RC elements indicate that the strains remained below or very close to the yield strain, meaning that the RC main structure remained elastic during the tests. This is corroborated by the fact that the damping fraction ξ obtained experimentally from free vibration tests conducted at the end of each seismic simulation remained constant (ξ = 1.8%). The maximum inter-story drifts IDI were 0.74% in the first story and 0.86% in the second one. These IDIs are slightly below and above, respectively, the value assumed in codes for the elastic limit of RC waffle-flat slab structures (0.75%). The corresponding ductility factors ${}_s\mu {}_i$ were ${}_s\mu {}_1=3.9$ and ${}_s\mu {}_2=4.0$ for the first and second stories respectively.

In contrast to the main structure, the WPDs underwent plastic deformations. Figure 18 and Figure 19 show the axial force versus axial displacement curves, N-d, measured on the metallic dampers installed in the first (Figure 18) and second story (Figure 19) in each seismic simulation. The energy dissipated by each damper at the end of the tests is obtained by integrating these N-d curves. Specifically, at the end of the tests the hysteretic energy dissipated by the damper of the first story was E_h1 = 4885 kNmm and that of the damper of the second story E_h2 = 3073 kNmm. The corresponding normalized ratios η_i = E_hi/(_sQ_yi × _sδ_yi) are η₁ = 33.4 and η₂ = 27.2 for the first and second stories. The values of _sQ_yi and _sδ_yi used for obtaining η_i were calculated with the equations developed in [34], as discussed in the next subsection.

5.5. Discussion

A nonlinear finite element model representing the test model without the dampers was developed using Engineer’s Studio software (V.1.07.02, Forum-8, London, UK) [35]. The numerical model was subjected to nonlinear static pushover analysis by applying an inverted triangle distribution of lateral forces. From this analysis, the lateral stiffness _fk_i and yield strength _fQ_yi of the main RC structure (without dampers) was obtained, giving _fk₁ = 3.6 kN/mm and _fQ_y1 = 20.6 kN in the first story, and _fk₂ = 4.1 kN/mm and _fQ_y2 = 22.4 kN in the second story. The lateral stiffness _sk_i and yield strength _sQ_yi provided by the dampers was calculated with the equations developed in [34] giving _sk₁ = 31.3 kN/mm and _sQ_y1 = 68.7 kN in the first story, and _sk₂ = 24.8 kN/mm and _sQ_y2 = 53 kN in the second story. Given these _sQ_yi’s, the actual yield shear force coefficients of the energy dissipation system ${}_s\alpha {}_i\left(={}_{s}Q{}_{yi}/{{\displaystyle \sum }}_{j=i}^N{m}_{j}g\right)$ corresponding to the lateral strengths _sQ_yi indicated above are _sα₁ = 0.63 in the first story and _sα₂ = 0.77 in the second story, and the distribution ${}_s\overline{\alpha }{}_i(={}_s\alpha {}_i/{}_s\alpha {}_1)$ is ${}_s\overline{\alpha }{}_1=1$ and ${}_s\overline{\alpha }{}_2=1.22$. Using Equation (16), the total yield shear force coefficients of the tested structure are α₁ = 0.70 in the first story and α₂ = 0.9 in the second story, and the distribution ${\overline{\alpha }}_i\left(={\alpha }_i/{\alpha }_1\right)$ is ${\overline{\alpha }}_1=1$ and ${\overline{\alpha }}_2=1.28$ for the first and second stories, respectively.

Meanwhile, idealizing the test model as a lumped mass system with two degrees of freedom (the horizontal displacement of each mass m₁ and m₂), and using the m_i and _fk_i indicated above, the vibration periods T_i and the corresponding vibration modes vectors φ_i of the RC main structure in elastic conditions were determined, giving T₁ = 0.15s, φ₁ = [1, 0.518] for the first mode and T₂ = 0.05s, φ₂ = [−0.3, 1] for the second one. With the above information, the ψ_mn,i parameters were calculated with Equation (22), giving ψ_m1,1 = 0.594 and ψ_m1,2 = 0.406 in the first mode, and ψ_m2,1 = 0.428 and ψ_m1,2 = 0.572 in the second mode. Further, using the input energy spectrum E_I-T of the accelerogram applied to the shaking table (Figure 17b) and applying the procedure explained at the end of Section 4.1, the inelastic hysteretic energy spectra E_h-T of the scaled accelerograms applied to the shaking table were calculated. Substituting the $E_{hn}^{SDOF}$’s provided by these spectra in Equation (17) and using the ψ_mn,i’s and other dynamic properties indicated above, the values of the optimum yield shear force coefficient distribution proposed in this study, Equations (17) and (22), were calculated giving ${\overline{\alpha }}_{prop,1}=1$ and ${\overline{\alpha }}_{prop,2}=1.22$. The values are very close to the actual yield-shear force coefficient distribution of the tested structure (i.e., ${\overline{\alpha }}_1=1$ and ${\overline{\alpha }}_2=1.28$). This explains why the cumulative plastic deformation ratios were very similar in the two stories (i.e., η₁ = 33.4 and η₂ = 27.2), as well as the ductility factors ${}_s\mu {}_i$ (i.e., ${}_s\mu {}_1=3.9$ and ${}_s\mu {}_2=4.0$), as shown in Section 5.4. This finding serves to validate, experimentally, the optimum yield-shear force coefficient distribution proposed in this study.

6. Conclusions

Past earthquakes have accentuated the importance of preventing a concentration of damage in multi-storey buildings, since it is a prevalent cause of collapse. Such an objective can be attained by using an appropriate distribution of lateral strength, one that can be expressed in terms of yield shear force coefficients. The present study focusses on structures with metallic dampers, proposing a new approach to obtain an optimum distribution of the yield shear force coefficients among stories, ${\overline{\alpha }}_{prop,i}$, that makes the damage–expressed in terms of cumulative plastic strain energy normalized by the product of yield strength and yield inter-story drift—approximately equal in all stories. The proposed optimum distribution is derived by applying modal analysis formulation and adopting two basic assumptions. One is that the damage spreads out evenly among the stories regardless of the level of plastic deformation. The second is that the ductility demand is approximately the same in all stories. These assumptions are supported by the results of past studies; the second one is further assessed in this study both numerically and experimentally. The proposed optimum distribution is validated through numerical simulations and shaking table tests. The new distribution if compared with the “exact” distribution ${\overline{\alpha }}_{exact,i}$ obtained from non-linear time history analyses, and with the distribution proposed by Benavent-Climent in previous studies, ${\overline{\alpha }}_{Benavent,i}$, leads us to these main conclusions:

(1)

${\overline{\alpha }}_{prop,i}$ is in good agreement with ${\overline{\alpha }}_{exact,i}$; the ratio ${\overline{\alpha }}_{prop,i}/{\overline{\alpha }}_{exact,i}$ ranges between 0.9 to 1.2 and the coefficient of variation (COV) is less than 0.10;

(2)

${\overline{\alpha }}_{prop,i}$ improves significantly upon the one proposed by Benavent-Climent. ${\overline{\alpha }}_{Benavent,i}$, particularly for the upper 1/3 of the height of the structure.

(3)

The interstory-drifts of the structures designed with ${\overline{\alpha }}_{prop,i}$ or with ${\overline{\alpha }}_{exact,i}$ are very similar.

(4)

Deviations of the yield-shear force coefficient distribution with respect to ${\overline{\alpha }}_{exact,i}$, inevitably causes deviations of η_i from the aspired even distribution of damage (i.e., η_i = η equal in all stories). Nonetheless, if such deviations of the yield-shear force coefficient distribution are less than about 20%: (i) the variations of η_i among stories are small, thus damage concentration becomes a minor problem, and (ii) the inter-story drifts scarcely deviate from those obtained with ${\overline{\alpha }}_{exact,i}.$

(5)

In structures having the optimum yield-shear force distribution, the ductility demand ${\overline{\mu }}_i$ is approximately the same for all stories.