Proceedings of the 9th International Conference on Structural Dynamics, EURODYN 2014 Porto, Portugal, 30 June - 2 July 2014 A. Cunha, E. Caetano, P. Ribeiro, G. Müller (eds.) ISSN: 2311-9020; ISBN: 978-972-752-165-4 Analytic – numerical model for walking person – footbridge structure interaction 1 Michèle Pfeil1, Natasha Amador 1, Roberto Pimentel2, Raimundo Vasconcelos3 Instituto COPPE, Univesidade Federal do Rio de Janeiro, POBox68506, 21941-972, Rio de Janeiro, Brasil 2 Department of Civil Eng., Universidade Federal da Paraíba, João Pessoa, 58051-900, Brasil 3 Department of Civil Eng., Universidade Federal do Amazonas, Manaus, Brasil email: [email protected], [email protected], [email protected], [email protected] ABSTRACT: In traditional dynamic analysis of footbridges the pedestrians are considered as moving time varying loads in vertical and horizontal directions. This approach is appropriate for most footbridges systems but for lightweight structures such as composite footbridges a high degree of human – structure interaction may be expected. This interaction mechanism, which has been reported in the literature, leads to a reduction in the footbridge response as compared to the response under the moving load model when there is one walking pedestrian in resonance with the structure. This paper presents an analytic – numerical model to address this feature in vertical vibrations by considering the coupling of a biodynamic model to a one degree of freedom system representing the structure. The walking person is simulated by a single degree of freedom mechanical model whose mass is set into vertical motion by heel´s elevations in forward steps. The model is applied to the Aberfeldy cable-stayed composite lively footbridge and the results are shown to be in good agreement with the experimental measurements performed under the passage of a single known pedestrian in resonance with the structure. KEY WORDS: footbridges, human-induced vibrations, pedestrian-structure dynamic interaction. 1 INTRODUCTION The current approach recommended in most codes of practice to estimate human-induced vibration amplitudes of vertical bending in footbridges is based on a moving load model whose magnitude is given by a Fourier series with 3 or 4 terms [1]. This procedure is appropriate for most situations but except for lightweight structures being crossed by pedestrians whose pacing frequency is close to a natural frequency of the system. In these cases it overestimates the maximum amplitudes [2] showing that the pedestrian – structure dynamic interaction has to be taken into account. References can be found in which this interaction is considered by coupling the bridge model to a moving springmass-damper (SMD) model to represent the vertical action of a walking person. The numerical model used by Fanning et al. [2] consisted in a moving mass connected to the upper node of a spring-damper element whose lower node was attached to a contact element through which the Fourier series load was applied to the numerical model of the bridge. The results focused on the vibration amplitudes given by the moving load model and the interaction model as compared to results from measurements of pedestrian crossings of Aberfeldy composite footbridge. Caprani et al. [3] addressed the same issue by developing a 2 degrees of freedom analytical-numerical model of the coupled system (bridge + pedestrian) and testing the influence of a range of parameters in the bridge response due to single pedestrian in addition to address crowd modelling. In a similar manner this paper presents a two degrees of freedom analytical-numerical model of the coupled system in which the walking person is simulated by a SMD model. The model is applied to the Aberfeldy cable-stayed composite lively footbridge. The theoretical results obtained from the moving load model and from the interaction model are compared to experimental results obtained by Pimentel [4] from two measurements of the structure under the passage of a single known pedestrian in resonance with the structure in order to excite the first and the second vibration modes. Correlations between the results focus on vibration amplitudes and also on the general aspect of time history responses. In relation to the parameters of the SMD model to represent a walking person in vertical direction there are few proposals regarding pedestrian – structure interaction studies. Kim et al. [5] proposed a two - degrees of freedom (2-DOF) biodynamic model whose parameters are, otherwise, applicable to standing people. For a single DOF model Caprani et al. [3] proposed values for the spring stiffness and damping coefficient within ranges taken from the human body biomechanic literature. In the present paper the single DOF model proposed by Silva and Pimentel [6] is used to simulate the walking person – structure interaction. This model was developed on the basis of vertical accelerations measurements of 20 tests subjects walking along a rigid surface and has the merit to allow application to each walking person as a function of its body mass and pacing frequency. ANALYTIC – NUMERICAL MODEL 2 2.1 Pedestrian walking on a rigid surface The walking person is represented by the single degree of freedom (SDOF) mechanical model with mass mp which is set into vertical motion by heels movements represented by displacement ur as illustrated in Fig. 1. The elastic and damping components of the contact force, Fep and Fap , between the pedestrian and the surface can be written as F t Fep Fep k p (u pr ur ) c p (u pr u r ) (1) 1079 Proceedings of the 9th International Conference on Structural Dynamics, EURODYN 2014 where u pr and u pr are respectively the vertical displacement and velocity of the subject center of mass while walking in a rigid surface; k p and c p are the pedestrian stiffness and damping respectively. The resultant force of eq. (1) is equal to the ground reaction force which can be represented as a Fourier series with 3 or 4 terms [1]: F (t ) M p g 1 .sen i.2 . f 3 i i 1 pp .t i (2) where ( M p g ) is the weight of the pedestrian, i and i are the dynamic load factor and the phase angle associated to each harmonic i of the load. mp upr kp displacement at the contact point k is U ek . The pedestrian model now displays a total vertical displacement equal to u p u pr , where u p is the displacement due solely to the pedestrian-structure interaction. The sum of the elastic and damping interaction forces at the contact point k depend on the relative displacement and velocity between the pedestrian and the structure and is written as: ur kp Figure 1. Analytical model for the interaction between a walking person and a rigid surface. k (3) The fundamental frequency fpp in eq. (2) is the walking frequency ranging between 1.5 to 2.4 Hz with a mean value equal to 2.0 Hz [1]. The magnitude of the dynamic load factors, particularly 1 , depend mainly on the pacing frequency f pp , but may display relevant differences among pedestrians and even for the same person depending for example on the shoes [7]. According to Zivanovic and Pavic [8] the most extensive research on the magnitude of dynamic load factor was conducted by Kerr [9] yielding the following expression for average values of 1 : (4) with 95% of the results falling in the range 1 0.321 . 2.2 The structure modal equation of motion is expressed as (5) where yi is the ith mode amplitude and mi , i and i are respectively the modal mass, damping ratio and frequency associated to mode i . In the moving load model the modal force Pi is determined from the applied force expressed by eqs. (2) and (4). 2.3 Pedestrian-structure interaction model Figure 2 illustrates the pedestrian model moving with constant speed along a flexible structure whose vertical 1080 e deformed structure m p up upr Fint (t ) (7) or as the following equation by taking into account eqs. (1), (3) and (6): m pup c p u p U ek k p u p U ek 0 (8) For the structure discretized in plane frame elements the vertical displacement and velocity at the contact point k can be expressed by U ek H.Ue and U ek H.U e (9) are the where H is the interpolation vector and U e and U e nodal displacement and velocity vectors of element e within which the contact point is located. The interaction force Fint t is applied to the structure model by means of the equivalent nodal force vector Fe of element e Fe HT Fint Structure model under human-induced moving load mi yi 2ii yi i2 yi Pi ur Uek The equation of motion of the pedestrian is then expressed as The equation of motion of the pedestrian model (Fig. 1) may then be expressed as 1 0.2649 f p3 1.3206 f p2 1.7597 f p 0.7613 cp Figure 2. Analytical model for the interaction between a walking person and a deformed structure. rigid surface m pupr F t (6) mp upr +up undeformed structure cp Fint t k p u p u pr ur U ek c p u p u pr ur U ek (10) By using the modal superposition method the vertical displacement at the contact point is expressed as U ek H Φ ei yi i ki yi (11) i where Φei contains the eigenvector amplitudes at the nodes of element e associated to mode i and ki is the magnitude of the modal shape at the contact point. The modal force is given by Pi ΦTei Fe (12) Proceedings of the 9th International Conference on Structural Dynamics, EURODYN 2014 F(t) Substituting the expression for Pi in equation (5) and taking into account eqs (1) and (2) leads to mi yi 2i i mi k2 c p y i i2 mi k2 k p yi k k p u p k c p u p k F t ut y X u p K k k p 2 (14) m 0 M 0 mp 2 m k2c p C k c p m k2k p k c p c p k k p k p F t F k 0 The equations are integrated by the Runge-Kutta method. The coupling terms in matrix K and C of eqs. (14) may also be written in terms of the pedestrian modal mass mp by substituting cp and kp by 2 p p m p and 2p m p respectively Figure 3 Single degree of freedom model used to develop the biodynamic model proposed by Silva & Pimentel [6] The model parameters mp, cp and kp were obtained by solving a system of three non-linear equations (eq. 16) which are the expressions of the accelerance frequency response function of the SDOF model for each the three harmonic of the forcing function eq.2 (i=1,3) i2 F (i ) (16) A(i ) k p i2 m p ji c p where j is square root of (-1) and F(i) and A(i) are the input values, being respectively the amplitudes of the harmonic components of the ground reaction force F(t) and of the measured acceleration at waist level for frequency i. Based on the observed correlations among the parameters of the model expressions were derived for the biodynamic parameters so as to take into account these correlations and allowing the parameters to be expressed as a function of the total body mass Mp and the pacing rate fpp. These expressions are given by: m p 97.082 0.275M p 37.518 f pp where p is the damping ratio and p the natural frequency of the pedestrian model. It can then be observed that the occurrence of the dynamic interaction is only to be expected when the pedestrian mass corresponds to a significant percentage of the modal mass of the structure. 3 SINGLE DOF BIODYNAMIC MODEL The biodynamic model proposed by Silva and Pimentel [6] is used in the present paper as the single DOF model representing the vertical action of a pedestrian walking along a rigid surface. It was developed on the basis of vertical acceleration measurements of the waist of 20 persons walking with their normal pacing rate on a straight path sufficiently long so as to ensure a regular movement. The model consists of a modal mass mp, a spring of elastic constant kp and a viscous damper whose constant is cp as shown in Fig. 3. These dynamic parameters of the model were sought in such a way that the measured accelerations at waist could be reproduced by the model expressed as follows. m put c put k put F t cp (13) Considering a single mode structural response, equations (8) and (13) form the coupled pedestrian-structure equations which may be written in matrix form as CX KX F MX mp kp (15) where ut is the displacement of the mass degree of freedom with respect to a fixed reference as shown in Fig. 3 and F(t) is given by eq. (2) with 1 expressed as Kerr´s expression eq. (4), 2=0.07 and 3=0.06. cp 4 29.041m0p.883 (kg ) (17a) ( Ns / m) (17b) k p 30351.744 50.261c p 0.035c 2p ( N / m) (17c) PRELIMINARY VALIDATION PEDESTRIAN-STRUCTURE MODEL OF THE INTERACTION Based on the observation made at the end of section 2.3 about the ratio between the modal masses of the pedestrian and of the bridge models an idealized simply supported 35m span steel-concrete composite footbridge was tested for the passage of a single pedestrian (simulated with the biodynamic model described in section 3) whose body mass Mp is equal to 75 kg. The pace frequency fpp was taken equal to 2Hz, coincident to the natural frequency of the structure corresponding to the first vertical bending mode. The ratio between the pedestrian and the structure modal masses is 0.25% and the structure damping ratio equal to 0.5%. The acceleration responses at the antinode of mode shape obtained with the moving load model (section 2.2) and with the pedestrian-structure interaction model (section 2.3) were exactly the same [10], as expected for this low ratio between modal masses. 5 STRUCTURE EXAMPLE AND MEASUREMENTS The footbridge used as example for pedestrian – structure dynamic interaction simulation is the known Aberfeldy cablestayed glass reinforced plastic (FRP) footbridge constructed in 1990 [11]. It is a three-span structure, having a main span of 1081 Proceedings of the 9th International Conference on Structural Dynamics, EURODYN 2014 Table 1. Measured dynamic properties Mode m (kg) f (Hz) Mode 1 S 2547 1.59 S+P 1.56 0.84 Mode 2 S 3330 1.92 S+P 1.88 0.94 S: isolated structure; S+P: structure + 80kg person Pimentel [4] also performed pedestrian tests on the cablestayed bridge in which a test subject of 80kg mass, departing from the south tower, walked with the help of a metronome at pace rates coincident with the natural frequencies of the coupled system: 1.56Hz and 1.88Hz, while acceleration measurements were taken at the antinodes of the corresponding mode shapes as shown in Figs. 5a and 5b respectively for the first and second vertical bending modes. acceleration (m/s²) 63m and two side spans of approximately 25m each. The towers were designed as A-shaped frames. Two groups of 10 cables connect the 2.12m wide deck to each tower. The results of the experimental tests campaign conducted by Pimentel [4] in July 1994 and June 1995 will be used in the present paper (other experimental assessments are reported in the literature [12] after reinforcement of the deck). It comprised a full modal survey - for the determination of the modal properties of the structure using ambient excitation, impulsive response and free-decay vibration - and an investigation of the bridge behavior under pedestrian-induced loads. From the ambient vibration tests the modal shapes of nine vibration modes and their corresponding natural frequencies were obtained. Fig. 4 illustrates the shapes of the first two vertical bending modes of the deck. 2,5 2,0 1,5 1,0 0,5 0,0 -0,5 -1,0 -1,5 -2,0 -2,5 acquisition ended while walking in the main span 25 35 45 55 time (s) 65 Figure 5a. Filtered acceleration time response from controlled pedestrian test in the footbridge excited at its first vertical bending mode [4]. Figure 4. First two measured vibration modes of the deck in vertical direction [4]. Equivalent viscous damping ratios were determined from the free-vibration decay part of the so called jumping tests in which the tests subjects were asked to bounce instead of jumping, as this was found to be an easy way to excite the structure at a pre-defined frequency [4]. The 80 kg mass test subject bounced in positions corresponding to the antinodes of the first and second mode shapes as illustrated in Fig. 4. There was no significant change in damping values along the freevibration decay records, which is an indication of linear behavior of the structure under normal usage. The results for damping ratios and frequencies of vibration obtained from the jumping tests are shown in Table 1 in the lines corresponding to the S+P (structure + person) indication. This table also shows the natural frequencies of the isolated structure as measured in the ambient vibration tests (see S lines) together with the modal mass m of each mode as obtained from a calibrated numerical model of the structure [4]. It can be noted that the person´s mass (80kg) is 3.1% the generalized mass of the first mode and 2.4% of that of the second mode. This is consistent with the observed lower frequencies of the coupled S+P system in relation to the isolated structure. 1082 Acceleration (m/s²) 1,5 1,0 0,5 0,0 -0,5 -1,0 -1,5 23 28 33 38 43 48 53 58 63 time (s) Figure 5b. Filtered acceleration time response from controlled pedestrian test in the footbridge excited at its second vertical bending mode [4]. 6 THEORETICAL RESULTS The moving load theoretical model (section 2.3) was applied to the cable-stayed footbridge with the properties shown in Table 2. Table 3 presents the parameters of the isolated bridge associated to the first two vibration modes and that of the biodynamic model of the walking person as adopted in the analysis of the dynamic interaction model (section 2.3). The properties of the biodynamic model were obtained with eqs (17) for Mp equal to 80kg. Proceedings of the 9th International Conference on Structural Dynamics, EURODYN 2014 4,0 Table 2. Dynamic properties of the coupled system used in the moving load theoretical model 0.84 0.94 f (Hz) 1.56 1.88 interaction model acceleration (m/s2) Mode m (kg) Mode 1 S+P 2627 Mode 2 S+P 3410 S+P: structure + 80kg person Table 3. Parameters of the structure (S) and of the biodynamic model (P) used in the interaction theoretical model for analyses of mode 1 and mode 2. m (kg) 1 S Mode 1 2547 P fpp=1.56Hz 60.5 2 S Mode 2 3330 P fpp=1.88Hz 48.6 P: Mp=80kg; fpp= pacing frequency. f (Hz) 1.59 2.67 1.92 2.65 0.84 91.6 0.94 78.0 moving load 3,0 2,0 1,0 0,0 -1,0 -2,0 -3,0 -4,0 25 35 45 55 time (s) 65 Figure 6. Acceleration time responses at the antinode of mode 1 of the footbridge under the passage of the pedestrian with pacing frequency equal to 1.56Hz, obtained with the moving load model and with the interaction model. 3,0 moving load acceleration (m/s2) 2,0 1,0 0,0 -1,0 -2,0 -3,0 25 35 45 55 time (s) 65 (a) 3,0 interaction 2,0 acceleration (m/s2) Although the damping ratios shown in Table 1 have been obtained from measurements on the coupled system these values were assigned to the isolated bridge (Table 3, S lines) in the interaction analyses presented in the present paper. This is justified by the relatively large spread between the vibration frequency of the bridge and the first natural frequency of standing people which is greater than 4Hz [13]. To confirm this hypothesis a theoretical-numerical analysis of the person – structure interaction was performed with a model similar to the one described in section 2 but with a biodynamic model of a standing person instead of a walking one [10]. This analysis showed that the 80 kg standing person simulated by a SDOF model with natural frequency equal to 4.9Hz and damping ratio equal to 37% [14] did not altered the structural damping ratio obtained numerically by free-vibration decay analysis. In the absence of force measurements on a rigid surface due to the tests subject, the dynamic load factors 1 (eq. 2) were initially taken as the mean values obtained from eq. (4): 0.224 and 0.360 respectively for modes 1 and 2. With these values the moving load and interaction models results were tested against the initial part of the experimental responses where amplitudes are low and both theoretical models give the same response eventually adjusting 1 for mode 1 to 0.270. Figures 6 and 7 show respectively the acceleration time responses at the antinodes of modes 1 and 2 of the footbridge under the passage of the pedestrian with controlled pacing frequency equal to the corresponding natural frequency of the coupled system obtained with two theoretical models: the moving load and the interaction model. It is seen in Fig. 6 that the two responses for mode 1 display the same overall shape as the experimental response (see Fig. 5a) and differ in terms of amplitudes, the moving load model yielding maximum amplitude 25% greater than that of the interaction model. 1,0 0,0 -1,0 -2,0 -3,0 25 35 45 55 65 time (s) (b) Figures 7. Acceleration time responses at the antinode of mode 2 of the footbridge under the passage of the pedestrian with pacing frequency equal to 1.88Hz, obtained (a) with the moving load model and (b) with the interaction model. The comparison between the theoretical responses obtained for mode 2 (see Figs. 7) and the experimental response shown in Fig.5b reveals that only the interaction model was able to reproduce the general aspect of the time response. With the moving load model the acceleration at the antinode of mode 2 drops to very low amplitudes when the pedestrian has just past the middle of the center span while the experimental and the interaction theoretical model results exhibit significant amplitudes at this same instant. Moreover the largest acceleration amplitudes are observed in the first part of the response obtained from the moving load model which is not the case of the experimental and interaction model responses. 1083 Proceedings of the 9th International Conference on Structural Dynamics, EURODYN 2014 7 COMPARISON BETWEEN THEORETICAL AND EXPERIMENTAL RESULTS Figure 8 shows the comparison for mode 1 test between the theoretical result obtained with the interaction model and the experimental result, the latter shown in terms of envelope points for the sake of clarity. It can be noticed a very good comparison until the experimental amplitude displayed a break from the expected trend of the response envelope indicated by the tangent lines in Fig.8. This could be due to the pedestrian losing step while perceiving footbridge vibrations as already observed by Zivanovic et al. [15]. In order to investigate this possibility a new analysis was performed in which the pace frequency was changed from 1.56 Hz to 1.57 Hz in the time interval estimated to correspond to the step losing as shown in Fig. 9. The results show that the model was able to reproduce the response by considering an idealized step losing. 3,0 2,5 1,5 0,5 0,0 -0,5 -1,0 -1,5 -2,5 23 28 1,88 1,88 2,5 1,88 2,0 1,88 1,88 1,5 1,88 1,0 1,88 1,88 0,5 1,88 0,0 1,88 -0,5 1,88 1,88 -1,0 1,88 -1,5 1,88 1,88 -2,0 1,88 -2,5 1,88 23 1,88 1,88 -2,0 55 time (s) Figure 8. Comparison between theoretical and experimental results for the first vibration mode. 3,0 43 48 53 58 63 interaction experimental acceleration ([m/s²) -1,0 45 38 Figure 10. Theoretical versus experimental results for the second vibration mode. 0,0 35 33 time (s) 1,0 25 fpp= 1.89 Hz 28 33 38 fpp= 1.89 Hz 43 48 53 58 63 time (s) Figure 11. Theoretical versus experimental results for the second vibration mode considering step losing. interaction model experimental 2,0 acceleration (m/s2) 1,0 -2,0 -3,0 8 CONCLUSIONS 1,0 0,0 -1,0 -2,0 fpp=1.57 Hz -3,0 25 35 45 55 time (s) Figure 9. Comparison between theoretical and experimental results for the first vibration mode considering step losing. The comparison between experimental and theoretical results for mode 2 is shown in Fig. 10. Again the theoretical results give larger amplitudes (in this case much larger) than the experimental ones. It can be noticed in the experimental response that step losing events may have taken place now for small vibration amplitudes and several times as indicated by color changing in the envelope points and tangent lines. A similar approach to that used in the case of the first mode response yielded the result shown in Fig. 11 for the second mode considering the step losing. As expected, the response is 1084 interaction experimental 2,0 interaction model experimental 2,0 acceleration (m/s2) very sensitive to minor changes in the step frequency. The correlation between theoretical and experimental results in Fig. 11 was not successful as for the first mode (Fig.9) but still the idealized step losing yielded a reasonable theoretical response as compared to the experimental one. acceleration ([m/s²) This is a clear demonstration of the occurrence of the structure – pedestrian dynamic interaction. An analytic-numerical model to address the dynamic interaction between a structure and a single walking person each represented by a single degree of freedom model was presented in this paper and applied to the all FRP cable-stayed Aberfeldy footbridge. The biodynamic model proposed by 29,84375 Silva and Pimentel [6] was used to obtain the parameters of 30,15625 30,46875 the biodynamic model. The developed equations made clear that the occurrence of the dynamic interaction is only to be expected when the pedestrian mass corresponds to a significant percentage of the modal mass of the structure, also depending on other parameters. For example, a preliminary test involving a simply supported steel-concrete composite footbridge whose modal mass was 0.25% of the modal mass of the pedestrian model that moved in resonant condition yielded equal responses from the traditional moving load model and the interaction model. The theoretical moving load and interaction models were applied to the Aberfeldy footbridge under the resonant Proceedings of the 9th International Conference on Structural Dynamics, EURODYN 2014 passage of a single pedestrian in each of two vibration modesthe first and the second vertical bending modes - in order to compare to experimental measurements [4]. In terms of maximum vibration amplitudes the moving load model led to values 30% greater than the interaction model. For the first mode the time responses obtained from both theoretical models displayed the same overall shape as the experimental response but for the second mode the comparison between the theoretical responses and the experimental one showed that only the interaction model was able to reproduce the general aspect of the time response. Observation of the experimental responses revealed sudden changes of the tangents to the envelope amplitude which may be due to pedestrian step losing. In terms of amplitudes over time the interaction model values for the first mode remained very close to the experimental ones until the pedestrian possibly loosed step for high amplitudes. For the second mode theoretical amplitudes grew beyond the experimental ones due to more than one step losing events. The correlation between experimental results and those from the interaction model considering idealized step losing was very successful for the first mode and reasonable for the second mode. In general the results point to the interaction model including the biodynamic model parameters proposed in [6] as an appropriate tool to simulate the bridge response under the passage of a pedestrian in resonance condition and as a conservative model to predict maximum amplitudes in this condition, yet not too conservative as the traditional moving load model. [11] W. Harvey, A reinforced plastic footbridge, Aberfeldy, UK, Structural Engineering International 4:229, 232, 1993. [12] J. Cadei and T. Stratford, The design, construction and in-service performance of the all-composite Aberfeldy footbridge, In Shenoi, A., Moy, S., & Holloway, L. (Eds.), Advanced Polymer Composites for Structural Applications in Construction. (pp. 445-453). ICE Publishing., 2002. [13] R.Sachse, A.Pavic, P. Reynolds, Human-structure dynamic interaction in civil engineering dynamics: a literature review, The Shock and Vibration Digest 35(351):3-18, 2003. [14] J.M.W.Brownjohn, Energy dissipation in one-way slabs with human participation, Proceedings of the Asian-Pacific Vibration Conference 99, vol.1:155-60, Singapore, 1999. [15] S.Zivanovic, A Pavic, P. Reynolds, Human-structure dynamic interaction in footbridges, Bridge Engineering 158 (BE4), 2005. ACKNOWLEDGMENTS The authors would like to acknowledge the brazilian agency for scientific and technological development Conselho Nacional de Desenvolvimento Científico e Tecnológico – CNPq for supporting this research. REFERENCES [1] H.Bachmann, W.Ammann, Vibration in structures induced by man and machines, Structural Engineering document n. 3, International Association for Bridge and Structure Engineering, Zurich, Switzerland, 1987. [2] P.Fanning, P. Archbold, A. Pavic, A novel interactive pedestrian load model for flexible footbridges, SEM annual conference, 2005. [3] C.C.Caprani. J. Keogh, P. Archbold, P. Fanning, Characteristic vertical response of a footbridge due to crowd loading, Proceedings of the Eurodyn 2011, Leuven, Belgium, 2011. [4] R.L.Pimentel, Vibrational performance of pedestrian bridges due to human-induced loads, Ph.D. thesis, University of Sheffield, UK,1997. [5] S-H. Kim, K-l Cho, M-S Choi and J-Y Lim, Development of human body model for the dynamic analysis of footbridges under pedestrian induced excitation, Steel Structures 8, pp333-345, 2008. [6] F.T. Silva, R.L. Pimentel, Biodynamic walking model for vibration serviceability of footbridges in vertical direction, Proceedings of the Eurodyn 2011, Leuven, Belgium, 2011. [7] W. Varela, Theoretical-experimental model to analyze human-induced vibrations on building floor slabs, D.Sc. Thesis, Instituto COPPE, Rio de Janeiro, Brazil, 2004. [8] S.Zivanovic, A Pavic, Quantification of dynamic excitation potential of pedestrian population crossing footbridges, Shock and Vibration 18(4):563-77, 2011. [9] S. Kerr, Human induced loading on staircases, Ph.D. thesis, University College London, London, UK, 1998. [10] N.A.Costa, Pedestrian-structure dynamic interaction in FRP composite footbridge, M.Sc. dissertation (in portuguese), Universidade Federal do Amazonas, Manaus, Brazil, 2013. 1085
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