Engineering Structures 29 (2007) 821–831www.elsevier.com/locate/engstruct Numerical and experimental study of vibration mitigation for highway light poles Luca Caracoglia a,∗ , Nicholas P. Jones b a Department of Civil and Environmental Engineering, Northeastern University, Boston, MA 02115, USA b Whiting School of Engineering, The Johns Hopkins University, Baltimore, MD 21218, USA Received 16 January 2006; received in revised form 29 June 2006; accepted 29 June 2006 Available online 14 August 2006 Abstract Highway light poles are slender structures usually characterized by low values of structural damping, a factor that can lead to large-amplitude vibration sometimes leading to collapse. This paper is motivated by a recent investigation, conducted to identify the reason for repeated failures, experienced by aluminum tapered light poles in the State of Illinois during a winter storm. The study combined numerical and experimental full-scale analysis of the structural system and its response to simulated external actions. It was observed that, despite the simple structural form, the definitive identification of the mechanism causing the oscillations was challenging due to inherent variability in the configuration as well as the paucity of environmental and response data. However, a plausible mechanism was identified, and a mitigation technique was proposed and evaluated for amplitude reduction. c 2006 Elsevier Ltd. All rights reserved. Keywords: Light poles; Dynamic loads; Vibration; Damping; Wind; “Frozen” precipitation; Galloping 1. Introduction Highway light poles are frequently subjected to environmental and wind loads. Despite the simple structural system and the availability of specific design tools (e.g., [1]), particular occurrences of unpredictable load configurations can lead to the failure of such systems or their sub-components. This paper is motivated by a research project managed by the Illinois Department of Transportation, initiated to understand the nature of some recent failures experienced on light pole structures in both serviceability and strength (collapse). One of the problems was the occurrence of large oscillation amplitudes experienced during a wind storm (Western Illinois) in tapered aluminum-alloy posts (Fig. 1(a) and (b)). These are typical slender systems, with reduced mass and low damping. As a result, approximately 140 units failed during this event. ∗ Corresponding address: Department of Civil and Environmental Engineering, Northeastern University, 400 Snell Engineering Center, 360 Huntington Avenue, Boston, MA 02115, USA. Tel.: +1 617 373 5186; fax: +1 617 373 4419. E-mail address:
[email protected] (L. Caracoglia). c 2006 Elsevier Ltd. All rights reserved. 0141-0296/$ - see front matter doi:10.1016/j.engstruct.2006.06.023 The structures had been designed in accordance with code recommendations. The occurrence of an unusual event or configuration was therefore considered as the main suspect, responsible for this incident. An extended investigation focused on potential mechanisms involving wind-induced vibration was conducted in order to consider the different aerodynamic and aeroelastic phenomena capable of leading to multiple system failures. The final objective of the research was not only the identification of the most plausible causes but also the evaluation of potential mitigation strategies to be directly applied to the existing structures and that could be incorporated in future applications. Therefore, an experimental campaign was designed and carried out to better address these issues, to identify the mechanical characteristics responsible for the interaction with the wind and to analyze potential solutions for the reduction of the amplification. 2. Description of the problem Aluminum-alloy tenon top poles (Fig. 1(a) and (b)) have been introduced on major highway interchanges in Illinois. The in many cases.P. usually responsible for the dynamic structural response. Illinois. higher flexibility results (E al = 70 GPa vs. Jones / Engineering Structures 29 (2007) 821–831 Fig. with the temperature constantly below zero degrees Celsius since the day before the collapse. The first component (10 min averaged wind speed.8]. The Gust Velocity Factor (GVF). The turbulence-induced zero-mean wind fluctuations in the along-wind. i. Uˆ (z). (1) with u ∗ the surface friction velocity and z 0 the roughness length. the potential combination of one or more of the above phenomena was considered. Installation dates of the poles in the Galesburg area were variable: most poles were installed in 2000. Poles were designed to withstand high winds in accordance with US specifications [1]. possibly initiated by the presence of a buildup of frozen precipitation on the surface of the pole or the luminaire. Caracoglia. u. site topography [6. is defined as the ratio between the maximum expected peak value of the velocity. These are also usually lowdamping systems. a fracture failure initiated at the handhole access in the pole itself. Estimates of u ∗ and z 0 are derivable for different terrain. IL [2].. and fatigue due to long-term cyclic loading (if applicable) were investigated. approximately 140 tapered aluminum light poles in western Illinois collapsed (at about the same time) at seven different interchanges on I74. These large oscillations induced failure either in the breakaway supports (anchorages to ground. in 1997 and others in 1999 [2]. (b) Example of collapsed units near Galesburg. The combination of lower mass with low damping increases susceptibility to wind-induced vibration. 3. tuned to the second mode of vibration. for neutrally stratified stable boundary-layer flows. Three main sources of excitation were identified: alongwind buffeting (dominated by first mode). G u (z) = Uˆ (z)/U¯ (z). as [3] U (z) = 2.5u ∗ ln (z/z 0 ) . For material capacity. During a winter storm on February 11. 1(a) and (b)). E steel = 205 GPa). are responsible for the dynamic response. located at approximately two-thirds of the height L with respect to the base. across-wind v and vertical w directions. 2003. on the contrary. the study focused attention on the response of the break-aways since a direct correspondence of load/stresses induced at the base of the pole was readily possible. (a). Wind environment Wind velocity is usually decomposed into a mean wind component. G u [1]. USA. and designed to mitigate dynamic amplifications due to K´arm´an vortex shedding [3–5]. and also equipped with canister vibration damper.822 L. and across-wind galloping [5]. vortex-induced vibration (at lock-in).e. most common reported cause) or. however. collapses involving the pole cross-section were more difficult to characterize and summarize due to an inherent variability in the failure configuration (Fig. 2003 (courtesy: Illinois Department of Transportation). At the perceived moment of the failure the recorded wind speed was about 22 m/s from the NW (steady winds with low turbulence). In addition.7] and exposure category [1. (c) Full-scale laboratory experiments. associated with long-term variability in the flow (10 min to 1 h) and a high-frequency component (order of seconds. turbulence). In the . N. main advantage consists in a lower mass (about three times lighter than steel). (2) The reference design value of wind speed (AASHTOIV. Heavy snow was also present. [1]) is defined as the three-second gust with fifty-year return period at z = 10 m (basic wind speed. BWS). as a function of the height with respect to ground z. 1. US-150 and US-34 in Galesburg. (1). and the averaged value of the along-wind component U¯ (z) through Eq. U ) can be represented. 13. Both maximum capacity of the material.7 m aluminum-alloy highway light pole. Two buffeting response characteristics were considered: the maximum peak equivalent displacement due to wind fluctuations in correspondence with the top of the pole (z = L. C Dn . Aerodynamic coefficients for the luminaries were estimated by comparison with equivalent rectangular sections and without specific wind tunnel tests. y. in particular compared to the maximum tensile capacity of the anchorages to ground (break-aways) determined from laboratory testing.x (z) and ψ j. (3) depending on the location z and the quantity that is monitored (for example. while the failure criteria are investigated in Table 1(b) .17 Hz. moment.x = 6. f 1. Caracoglia.9 Hz for larger M L .x = 0.49 Hz. Tapered light poles (13. where the first-mode buffeting response for a 13.61 Hz. y or C “equivalent combination” if both x–z and y–z loads are considered). The dynamics characteristics of the poles were represented by considering symmetry in the along-wind. a CWS of 28–30 m/s. ymax (with Y¯ = 0). Dead load of the pole with the potential presence of precipitation accumulation was negligible. f 2. Although the asymmetry in the luminaire fixture can produce twist and bending. Analytically-derived frequency and mode 823 shapes were also compared with finite-element simulations. Different luminarie models were present with mass M L estimated between 18 kg and 27 kg. x.000 [3].L. This value is located in the proximity of the range where the critical transition from laminar to turbulent flow occurs. For this area (exp. δ D (between 5 and 7 mm). i. directions.35 Hz. The first mode frequency is located in the neighborhood of 1.61 Hz.. linearly variable with height.4 m/s and z 0 = 0.Q (L) (Q is labeled as x.7 m) with circular cross section and variable wall thickness.e. gust effect factors G Q . with the same frequency but perpendicular orientation. An example of the simulation results is included in Table 1(a) and (b). the correspondence was very good [15]. First-mode dominated response was considered. stress at a particular cross section). All quantities are computed for z = L. but compatible with the results for smooth cylinders at Re = 250. f 2. The outer diameter. Similar studies are available for thicker units. total modal damping ζ1. Strength failures associated with buffeting Buffeting analyses were based upon the closed-form method proposed in Ref.1 independent of the Reynolds Number Re was considered in this section. 1(b)) are active and a second. and including fluctuations in the wind energy content at different frequencies (background response). The generic peak response effect R(z) at a specific location z. dynamic amplification due to the interaction with the structure (resonant response) and partial loss of correlation of the pressures due to the span-wise extension of the system. A constant value of C D0 = 1.x = 6. In Table 1(a) the effect associated with the variation of the surface roughness is analyzed.x = 1. 4. in which two inactive bolts lie on the neutral axis.x = 7. was employed. n = 1 luminaire) were specified in accordance with the values indicated by AASHTO-IV. torsional components were neglected since the eccentricity was considered as secondary in the analyses.and second-mode frequency for δ D = 7 mm and luminaire mass M L equal to 0. are: f 1. Jones / Engineering Structures 29 (2007) 821–831 analyses the Characteristic Wind Speed (CWS).52 Hz. xmax and ystd . [10] for the study of masts and monotubular steel towers due to fluctuating wind loading under stationary ˆ wind conditions. and estimated as 1.5%. For thinner poles (δ D = 5 mm) the frequencies are: f 1. Normalized aerodynamic force coefficients (drag. maximum along. effects will be investigated in the following sections.x = 1. M L = 13 kg to that used in the laboratory experiments (Section 9 and Fig. ψ j.e.x = 1.y (z).Q (L) = X¯ (L) + xeq. C Ln with n = 0 pole.02 m. The quantity M L = 18 kg corresponds to the luminaire mass of the models installed in the field.. and acrosswind. is referred to its corresponding mean value ¯ R(z) through an appropriate and unique Gust Effect Factor (GEF) [11]. Frequencies and modes were derived in closed form by simulating the system as a cantilever beam with tapering cross section and the representation of the structural mode shapes in terms of Bessel functions [12–14].x = 1. root-mean-square and maximum dynamic displacements in the two orthogonal planes xstd . f 1.Q .0 Hz. as later confirmed by the experimental results (Section 9). f 1. Two shear/bending moment mechanisms were considered: a first mechanism in which all the four bolts (Fig.14 Hz. Structural modal damping of the poles (percentage with respect to critical) was assumed as ζ1.18 Hz. a set of j independent mode shapes.x = 6.10 Hz. 1(c)). f 1.x = 7. with L height of the pole) and the maximum buffeting-induced overturning moment and shear forces at the base. ˆ ¯ G R (z) = R(z)/ R.x = 6. z 0 ) is summarized.P. f 2.31 Hz. These tables show the maximum peak displacement at the top of the pole X eq.49 Hz. more stringent mechanism. f 2. The Solari wind spectrum was used [17] with meteorological and topography conditions at the time of failure simulated through u ∗ = 1.x = 1.24 Hz. For comparison purposes of CWS with BWS.x = 1. characterized by reduced values of drag. [10]). N. the GVF was computed for exposure category C “open terrain with scattered obstructions”. The presence of both along-wind and across-wind buffeting was simulated by independently computing the first mode along-wind and across-wind response (and equivalent dynamic loading) and combining the two effects by means of an appropriate combination rule (non-simultaneous occurrence of the peaks in the x–z and y–z planes. Typical values of the first. tending to 0. C) a BWS of 40 m/s is prescribed [1].or across-wind displacement.41 [9]. was the same for all cases: 152 mm at the top of the structure and 254 mm at the base. compatible with similar metal units [16]. lift. i. The failure criterion (wind condition necessary to exceed the tensile capacity of anchorages) was selected in accordance with a collapse limit state methodology and the most vulnerable anchorage in both shear and tension. had been installed. 13 kg and 18 kg. as a conservative value with respect to the code. f 2. shear. referred to 10 min averages.31 < G u < 1.7 m pole with δ D = 5 mm and different wind characteristics (u ∗ . respectively. f 2. A full description of the procedure can be found in [15]. The mechanical characteristics of the material were derived from AASTHO-IV. also considering variability of C D0 and C D1 .C (L) can vary in the range 0. (moment.60 m (Table 1(a)). 6.824 L. 20]. U (10) = 22 m/s. the moment and shear forces recorded at the base are approximately two to three times lower than the values necessary to induce failure in the break-aways. and dynamic interaction can be excluded due to evident effectiveness of the canister vibration damper (Section 2) tuned to the second mode.P. (b) U (10) = 22 m/s. In particular. In the tables. usually beneficial in the redistribution of stresses during the lifetime of the structure. depending on the failure mechanism and considering the effective action of a limited number of anchorages (two) due to wind direction. third to fourth mode). For an average wind and turbulence scenario (z 0 = 0. load cycles). such as the statistical distribution of the wind speed and the reference fatigue curve of the material (maximum differential stress vs. suggests a minimum value of 3. f = 22 Hz (i.50 < xeq. u ∗ = 1.x + ζ1. along-wind turbulence intensity Iu = 0.45– 0. Simulations indicated that the fatigue life of the breakaways could be conservatively estimated as about eight years. 5. In general. Jones / Engineering Structures 29 (2007) 821–831 Table 1 First-mode gust buffeting response for a 13. Lockin from vortex shedding was not responsible for failures. U (10) = 22 m/s and Dref = 203 mm (average value of outer diameter). V0C max and derived as equivalent static loads [10]). this fact was observed through a parametric investigation conducted by varying the material reduction factor . The first-mode total damping in the along-wind direction. (a) u ∗ = 1.Aerod (structural and aerodynamic components) is significant and mainly related to the beneficial effect of the luminaire: 4%–5% at U (10) = 22 m/s. variable z 0 and u ∗ 1st (.7 m aluminum-alloy pole with δ D = 5 mm and ζ1. shear at the base indicated as M0C max . Fatigue failures derived from long-term exposure to buffeting or lock-in conditions Fatigue failures due to wind buffeting were also investigated by means of the analytical method recently proposed in [18] for the study of cylindrical vertical structures affected by wind loading.4 m/s and variable z 0 . and a mean deflection equal to 0.26 m under steady wind for δ D = 5 mm. Caracoglia. Wind directionality effects. as integrated by recent studies [19. Winds with constant direction were considered with 20% probability of absence of wind [18].x. Influence of vortex shedding on the response during collapse Vortex shedding in the across-wind direction was neglected since it usually involves higher modes. dashed–dotted line for mechanism two) were computed. including the effect of the potential presence of geometric tolerance in the installation of the anchorages to ground.02 m. Fatigue life for the aluminum cross-section at the base was also estimated as about seven to eight years.. requiring much higher winds or physically unreasonable turbulence intensity levels. these predictions are extremely sensitive to the input parameters. corresponding to a dynamic oscillation with 0.e.5%.60 m peak-to-peak amplitude (Table 1(b)). (4) suggests that for S = 0. were neglected. evidenced by the observed large-amplitude oscillations and the fact that the Strouhal relationship.–) failure mechanisms for variable geometric tolerance. the wind conditions associated with both failure mechanisms (dashed line for mechanism one. by taking into account different reduction factors in the definition of maximum fatigue capacity for aluminum.x.4%.Struct = 0. Despite the moderate amplitudes.2.16) X eq.4 m/s. The influence of the drag coefficient in the critical range of Re was also carefully analyzed. It was concluded that the wind characteristics or the extremely large vibration necessary to exceed the tensile capacity of the anchor bolts were not consistent with the meteorological conditions at the time of the event.C (L) < 0. S = f Dref /U.-) and 2nd (– . estimates can differ by one order of magnitude due to the nonlinear nature of the fatigue curves. ζ1. A further parametric investigation.tot = ζ1. N. The quantity C L0 + C D .P.y = 0.tot > 0) and unstable dynamic oscillation (ζ1. the latter value also being suggested by the AASHTO-IV interpolating equation for a wind speed compatible with the meteorological conditions of the event. In particular. Specifically. This suggests that slurry/freezing precipitation could have possibly adhered to the surface of the pole. as UCR / ( f 1 Dref ) = −4 C L0 + CD −1 −1 2 2π m 1. turbulence fluctuation effects on Eq.000 and for cylinders with dimensionless roughness coefficient of the order of 10−4 to 10−3 (metal surface). a combination of the drag coefficient per unit length C D and the first derivative of the lift coefficient with respect to the angle of attack C L0 = ∂C L /∂α. The subsequent application of the closed-form method proposed in [21] to this specific case suggested a life expectation variable between two and seven years for the aluminum pole base cross-section. Eq.eq ζ1. separately. a concurrent collapse of systems located in the same area with different failure characteristics (and service ages) exclusively associated with fatigue was considered as statistically improbable. (5) The vanishing of the total damping in Eq.0. Results are summarized in Table 2.tot < 0).y ρ Dref . m 1. As an example. in the critical range of Re. although a penalizing factor relating stress variation to wind speed was employed due to the limited information and data. the critical wind speed derived for a 5% ice thickness (15–18 m/s — cases (ii) and (iii)) is low for the pole that was considered relative to the value at the time of collapse (22 m/s). Linearized analysis and a quasi-steady approach can be adopted since typical frequencies of the poles were much lower than those associated with vortex shedding for U = 22 m/s [5].y is the structural modal damping. Lift coefficients. it was found that for Re = 250.7 m pole with M L = 18 kg. also considering the fact that most units were installed between 1998 and 2000. Attention was devoted to the study of the drag coefficient.85 [3].55 and 0. (6) where f 1. 7. ρ is the air density. the presence of the impact canister damper and the indication of large-amplitude vibration at the time of the incident. can supercool at low temperatures when falling from the upper levels of the atmosphere due to barometric pressure gradient.6L).y . In a recent study [21] it was shown that fatigue life of steel poles can be as low as one year under special wind exposure conditions and due to the effects of drag reduction in the critical Re region. The necessary condition for across-wind first-mode galloping is ζ1. C D0 can vary between 0. Although symmetric circular sections are immune from this type of phenomenon. Analyses included variable geometry and wind characteristics. Dref is a reference dimension of the cross section (at z = 0. although the simulated deposit of frozen precipitation seems . (5) can be translated into a critical wind velocity for galloping onset for a single degree of freedom system [3. for a 13. C L0 and C L0 ’ were derived in accordance with the experimental data in [22] for a circular cross section with ice thickness variable from 3% to 5% with respect to the reference diameter of the body and a coverage area of approximately 120◦ . the meteorological situation corresponding to the investigated event suggests that the potential presence of frozen precipitation on the surface of the pole could not be excluded.L.eq is an equivalent first-mode mass per unit length (including the presence of the luminaire). For the basic value of C D0 = 1.y. as indicated in [23]. Since ζ1. N. (6) were neglected.y. Jones / Engineering Structures 29 (2007) 821–831 (fatigue limit states) either applied to the aluminum crosssection (pole base) or to the break-aways tensile strength to simulate the effects associated with the different fatigue categories (severity of the loading).5%. Fatigue issues related to vortex shedding were also excluded due to the extremely low value of lock-in velocity range corresponding to the first mode. In addition. ζ1. In contrast.y + ζ1. for an equivalent 120◦ coverage area from the top to the base (case (i) in Table 2).22]. buffetingrelated fatigue did not seem compatible with the multiple simultaneous occurrences experienced in this particular event. Moreover.73 according to [25]. experimental studies [24] report that some luminaire shapes are potentially prone to gallopingtype oscillation (in the absence of frozen precipitation). Since the analysis of a collection of different experimental tests [26] indicated considerable variability of C D0 . and 0. Super cool water (liquid below freezing temperature) freezes almost immediately when reaching another surface characterized by a significant temperature difference with respect to the air. precipitation such as rain or snow close to freezing temperature.y.y. 5]. variable wind profile with z 0 = 0. C D0 . (6) was cast into a more general formulation for the analysis of first-mode galloping oscillations of poles and vertical masts [10. must be negative (Glauert–Den Hartog Criterion). triggered by a decrement of the aerodynamic damping component and involving energy transfer from the wind flow. Across-wind galloping instability Galloping is an aeroelastic instability phenomenon involving large-amplitude oscillation in the across-wind direction [3. The mean wind velocity UCR was referenced to z = 10 m since the presence of 825 a boundary-layer profile was considered. by incorporating modal quantities and simulating the presence of frozen precipitation on the whole surface of the pole and in correspondence with the top of the luminaire. Eq.Aerod is a function of the wind velocity.0 the galloping critical velocity UCR at z = 10 becomes extremely large when the surface of the pole is covered by 3% ice. Caracoglia. (5) can be interpreted as the boundary between stable (ζ1.Aerod = 0. a sensitivity investigation was conducted by evaluating distinct limiting conditions with respect to a reference value C D0 = 1.5]. with ζ1.y is the across-wind first-mode frequency of the system. Fatigue-life predictions larger than seven to eight years seem compatible with previous results derived for steel poles through the method utilized in [18].02 m and assuming modal damping and mass distribution in accordance with mode shapes ψ1.tot = ζ1. for example. UCR .4 1. seemed consistent with the wind and environmental conditions. based upon a statistical method for the numerical simulation of the failure probability ( p f ) due to galloping induced by simulated precipitation deposit.826 L. Vibration mitigation The recommendations for vibration reduction were proposed by recalling the definition of the instability criterion for windinduced vibration of cylinders and masts.0 −2. Jones / Engineering Structures 29 (2007) 821–831 Table 2 First-mode galloping critical velocity. galloping due to frozen precipitation exclusively concentrated on the luminaire fixture was not a plausible explanation.0 is suggested in [27]. 2.3 −3. no C D0 .0 0. low C D0 .3 −5. leading to an increase of galloping instability threshold.0 1.0 1.0 −2.5 0. small luminaire fixture) were the most vulnerable to this type of aeroelastically enhanced phenomenon.5 −1.0.0 0. which would lead to extremely low velocity values (case (v)). partial coverage of freezing rain.4 1. this problem was typical of aluminum units since it was also shown that for equivalent steel poles the same behavior was associated with much higher and physically unreasonable wind velocities. This condition.9 0. including. (6)).4 1.R /Dref (normalized root-mean-square response). whereas the galloping onset corresponds to the asymptotic behavior at UCR . lock-in from vortex shedding). distributed on a selected portion of the pole. further experimental and wind tunnel studies would be desirable to better ascertain the aerodynamic coefficients of the complex luminaire crosssection and to exclude or confirm its influence on the event. . L = 13. according to which Fig. In 1 A recent development of this research.0 −1.. This quantity is a measure of susceptibility to aerodynamic instability (typically. for a 13.4 0. In any case. suggested that the original threshold indicated in this paper may be replaced by N = 3. From Table 2 and the exclusion of other factors (Sections 4– 6). The positive quantity N varies from case to case. (conserv.4 0.0 −5.0 0. at z = 10 m. C L . 8. “ice” on lum. This observation also excludes the direct influence of the drag reduction and Re.0 1.4 was computed for the analyzed case. i. Caracoglia.4 1. The simultaneous or alternate vanishing of C D0 and C D1 led to estimates closer to the target value. similar to cases (ii) and (iii) in Table 2.. C D1 7 7 5 7 7 7 7 7 1.4 0. 2 and indicated by U R and σ0.e.0 0.9 −1. perhaps combined with unstable luminaire shape and affected by meteorological conditions. the equivalent first-mode dimensionless Scruton Number.7 m pole with M L = 18 kg as a function of wall thickness. since the critical speed derived from the simulations was either extremely high (case (iv) in Table 2) or physically unreasonable (cases (vi) and (vii)). and variable C D . low C D0 Simulated “ice” on lum. aerodynamic damping devices [29].4 0.0 0. an aeroelastic phenomenon initiated by the build-up of freezing rain and characterized by large values of the resonant portion of the dynamic response power spectral density [10]. traffic signals and luminaries susceptible to wind excitation [20.. but this circumstance does not seem physically possible (high surface roughness of the poles. Schematic representation of galloping threshold UCR and aeroelastically enhanced phenomenon as a function of normalized root-mean-square across-wind response.5% damping. (7) must satisfy the condition K S > N (see Eq. equivalent to buffeting prior to galloping. for a reference diameter (0.28].4 1. however.0 0. Different damping devices have been proposed for light poles.0 0. Furthermore. equivalent mass corresponding to first-mode vibration and 0. A subsequent analysis [23] demonstrated that this class of aluminum poles (thin wall.). δ D . −1 2 K S = m 1.2 m). is schematically depicted in Fig.. C D1 Simulated “ice” on lum. low C D0 Simulated “ice” on lum.2 and 1. possibly.eq ζ1. N.y ρ Dref .0 69 18 15 110 14 208 300 17 extremely penalizing and incompatible with the meteorological conditions during the investigated event.0 1.P.7 m. while a K S between 1. a value of 4. for “masts with roughness or longitudinal attachments in the critical Reynolds Number regime”.0 0. and luminaries are bluff bodies with large flow separation). (conservative) “Ice” on lum. However.1 An increment of structural damping was strongly indicated as a possible solution. TMD/impact dampers [30] on long horizontal arms of traffic lights or impact ball dampers on posts to reduce K´arm´an-vortex-induced oscillation [31]. Moreover. approximately equivalent to 10−4 < p f < 10−3 (yearly value) [23]. it was concluded that the most plausible multiple-failure cause could be related to across-wind galloping instability or. C L0 (i) (ii) (iii) (iv) (v) (vi) (vii) (viii) Case δD (mm) C D0 0 C L0 C D1 0 C L1 UCR (m/s) 3% simulated “ice” on pole 5% simulated “ice” on pole 5% simulated “ice” on pole Sim. An appropriately designed device (l = 3. and an equivalent viscous damping coefficient. chain dampers are not effective at very low amplitudes or. for example.8 m) were digitally bandpass filtered to represent the response associated with the fundamental mode without (a) and with (b) chain damper.e. with estimated modal damping equal to 7. the component related to soil-structure interaction at the base/foundation of the pole.. Closelyspaced but distinct frequencies (1%–2% difference) were also .22 g/l frequency of the pendulum. The testing involved different configurations and combinations: with and without the luminaire. full or no impact occurs. directly in the field [16]. corresponding to the case without luminaire. ψ j. for example. one in-plane accelerometer (AC3) located at z = 0. chain length l. 9. USA) to study the dynamic response characteristics of aluminum tapered poles and validate the proposed solution. The corresponding initial deflection at the top of the pole. x–z plane of the luminaire in Fig. 827 A vertical orientation was preferred due to the versatility of the apparatus and the necessity of testing the proposed mitigation device. while beyond a critical value of frequency ratio (depending on d/x0 ). (8) also shows that the damping is amplitude dependent. University of Illinois at Urbana-Champaign. ceq . the interaction between canister damper and the new device was considered. most probably related to nonhomogeneity and asymmetries of the pole in the x–z and y–z planes. Unfortunately. i. high frequency [32. ceq /lmω optimal = 4π −2 d/x0 . Fig. 3(a) and (b) show the time histories associated with the freevibration test no.5% for d/x0 = 0. Experimental investigation of the proposed damper performance Due to the inherent uncertainty associated with the definition of the impact damper performance an experimental campaign was conducted in the NSEL (Newmark Structural Engineering Laboratory.e.7 m tapered straight pole with δ D = 5 mm. 1(c)).06 < x0. y–z plane and orthogonal to x–z) single-channel accelerometers at the top of the unit. For ω/ωCh > 3 but below a critical value the damping group becomes approximately constant (optimal) and independent of the excitation frequency. The quantity x0 can be interpreted as the amplitude associated with the firstmode generalized coordinate. Also. The range of initial conditions 0. x0. anchored to the floor of the laboratory (z = 4. Climate and temperature (such as. The specimen was anchored to a heavy concrete block. During the free-vibration tests the structure was quasistatically pulled backward in the x (in-plane) direction through a hydraulic actuator. free-vibration tests were performed in order to extract the frequency and the damping characteristics of the investigated pole. This condition can be estimated as [32]. direct experimental simulation of these effects was not possible.x and ψ j.13 m (the minus sign denotes deflection due to backward pulling). through an opening in the strong floor. the reported conditions at the moment of collapse) may have detrimental effects on such values. a 13. a luminaire with M L = 13 kg was used.0% for d/x0 = 1. for moderate oscillation amplitudes. although it can be mitigated through a rubber covered chain without significant reduction in the impact restitution coefficient [32]. This phenomenon is not associated with the inherent damping of the pole but with the presence of a “false beating” effect.0 (depending on amplitude). i. √ as a function of the frequency ratio ω/ωCh .L < 0.. commonly used on tall masts or slender truss-type towers and sometimes on shorter vertical systems.L . referenced to an existing reaction wall.33] derived for a device suspended inside a container attached to a vibrating structure. Noise is a disadvantage of the approach. The laboratory tests confirm that aluminum poles are structures where the inherent damping component is low and that the introduction of a damper was desirable to reduce the sensitivity to aeroelastic excitation.L = −0. The design of such devices was performed by employing a series of optimal damping curves [32. a first-mode Scruton Number K S > 7. 1(c) shows some of the installation and experimental stages. 3(a)) an initial oscillation decay is followed by an increment at t = 25 s and subsequent second decrement. i. Vibration reduction is given in terms of a normalized damping group.46 m with respect to the pole base) and released from this initial condition. 01BN040. The best performance is achieved when two impacts per cycle occur. Different operational regimes can be observed.0 (Eq.56 L = 7. after the removal of the canister damper and x0.P. N. Fig. Prior to the installation of the device. observable even in the absence of the eccentric luminaire. Experimental studies conducted on steel mono-tubular masts. (8) Eq. ceq (lmω)−1 . Original data (in-plane displacement at z = 11. (8)) and 3. Caracoglia. was derived from the available instrumentation. An imperfect planarity of the x and y modes. with ωCh = 1.30 m was considered to study the amplitude-dependent response of the chain.05 m. Instrumentation included in-plane (AC1.y .6 m. The pole was originally equipped with canister vibration damper (Section 2) tuned to the second mode. in both the absence and the presence of the chain impact damper. At large oscillation amplitudes the number of impacts usually increases and becomes irregular. In the absence of chain damper (Fig. subject to forced vibration at angular frequency ω and constant amplitude x0 .L. similar to those involved in the multiple collapses. m = 1. mass per unit length m. Jones / Engineering Structures 29 (2007) 821–831 this study a simple yet effective and inexpensive method for the reduction of wind-induced vibration on vertical structures was considered: a chain damper. 1(c)) and out-of-plane (AC2. located in the basement of the laboratory. in which partial. suggest that structural damping of such systems is variable and can be a function of several parameters including..34]. an equivalent dynamic displacement at the top of the pole.48 kg/m and ω/ωCh ∼ = 3) was employed. damping drastically diminishes (no impacts). As an example. close to the anti-node of the second mode and a displacement transducer at z = 11.5.e. is responsible for a two-dimensional response.8 m in the x direction (Fig. equivalently. However. In Fig. the combination of both devices is desirable. (a) “bare pole”. usually observed during the tests with chain damper. canister damper removed). N. 3. Regular impacts (two per cycle) were noticed in most cases for a relatively wide range of amplitudes. which can be dissipated through the canister damper. time history recorded at z = 11. measured) is ζ1. Acceleration time-histories associated with the fundamental mode (after band-pass filtering) are depicted. during which the effectiveness of the chain is drastically reduced (“whirling motion”). due to the effectiveness of the passive control device. In these cases damping was evaluated from the residual oscillation in the out-of-plane direction extracted from AC2 (ydirection).417 Hz. recorded. Fig.13 m before release). from in-plane displacement time history at .8 m (x0. respectively corresponding to Fig. The chain damper alone is not effective when the response is dominated by second or higher modes. therefore. Fig.36]) were employed but the extraction of damping ratios through mode decomposition was still challenging and influenced by the closeness of the modes.L = −0. Amplitude rapidly decreases after a few cycles and the two-dimensional response of the pole is reduced.x = 1. M L = 0 and canister damper removed). 3(b) the two-dimensional motion is suppressed due to the high levels of damping provided by the hanging chain. since the latter is ineffective at low amplitudes. The coupling of the chain with the existing passive device was also analyzed. Alternative time-frequency multi-scale techniques (e.P. measured). during the free-decay evolution the pole typically undergoes elliptical or circular trajectories with moderate oscillation amplitudes for a limited duration. No significant first-mode response differences were observed before and after removing the canister damper. 4(a)) and the interaction between x and y response components. Jones / Engineering Structures 29 (2007) 821–831 Fig. Nevertheless.y = 0..7 m pole free-vibration tests (M L = 0.2%. In-plane (AC1) and out-of-plane (AC2) first-mode filtered accelerations (13. 4(a) and (b). 3(a) with the addition of the chain damper. one of the perceived advantages is related to the promotion of energy transfer from the fundamental to higher modes. and restricted to the interval 5 < t < 20 s.24 m before release and in the presence of chain damper. in consideration of its location and the presence of amplitude-dependent critical frequency ratios [32].418 Hz. Example of 13. x0. Accurate damping estimates in the absence of the chain were particularly difficult due to the presence of large-amplitude twodimensional motion (Fig. The separation of the two modes by signal filtering was not possible. 5 presents an example of a first-mode damping estimate for M L = 0 ( f 1.L = −0. also confirming the predictions (Section 8).13 m before release. 3(b) is related to the same configuration as in Fig.L = −0. (a) “bare pole”. A typical value derived for M L = 0 and for the out-of-plane mode ( f 1. 3(a) and (b).7 m pole. (b) with chain damper.828 L. [35. (b) with chain damper (band-pass filtered signal). 4.g. with x0. Caracoglia. An example of a two-dimensional trajectory in the absence and presence of chain damper is presented in Fig.y = 1. Fig. a.9 1.L = −0. a “kink” is evident at t = 13 s.8 m. Time intervals for the calculation of ζ1 were selected in order to be representative of large (x0. Jones / Engineering Structures 29 (2007) 821–831 Table 3 First-mode structural damping (ζ1. chain (test01BNC081) M L = 13 kg. note 1) M L = 13 kg. experiments suggest an equivalent ζ1.L < 0. Other techniques for the derivation of ζ1 were analyzed such as a method based upon a direct nonlinear regression of either acceleration or displacement records.P.12 3. (b) damping ratio estimates (least squares) through envelope curve. as a consequence. without chain. 0. chain (test01BC040) Note 1: ζ1. This method is sensitive to the time interval (between 8 and 20 s in Fig. In the table.a. From least-squares linear regression in the semi-logarithmic plane of A(t) (timevarying amplitude derived through Hilbert transform [37]) an equivalent viscous damping corresponding to the selected mode can be derived.a.2%. 0.L (m) ζ1 (%) M L = 0.1 −0. approximately ten times larger than the original value. N.158 (1.829 L. associated with the limit beyond which no impacts are recorded. for amplitudes A = 0.8).38 m peak to peak at z = 11. Example of damping assessment from free-vibration tests (13. note 1) M L = 13 kg. Results confirm low values of first-mode structural damping component. The equivalent estimated curve corresponding to linear damping is also indicated in Fig. chain (test01BC080) 1. NO chain (test01BC080. 0. This effect was interpreted by the fact that the eccentric mass at the top of the pole promotes elliptic trajectories and. chain (test01BNC040) M L = 29 Lb. 5(b)) and inactive chain (t > 13 s). Equivalent viscous damping was derived for the classes of large (Table 3(a)) and moderate (Table 3(b)) oscillation amplitudes.12 2.a. Differences can be attributed to the fact that residual oscillation regimes (inactive chain) are influenced by the initial conditions and the decay trajectories linked to the coupling in the x and y directions.412 n. A(t). It is worth recalling that such values were conservatively derived as shown . Fig.418 (1. NO chain (test01BC040.24 4.7 m aluminum alloy highway pole for large (a) and moderate (b) oscillation amplitudes (a) Case f (Hz) (note 2) x0.1 −0.x ) of a 13. An overall reduction of damping is observed in the presence of the luminaire at both moderate and large amplitudes with respect to the case with M L = 0. Although the procedure allowed for the derivation of damping as a function of amplitude. In general. between 0.4 (b) M L = 0.2 m) and moderate (x0.1 m) vibration amplitudes. In these cases amplitude dependence is not applicable (n.158) 1.a. In the figure.2 −0. the computation suggests ζ1. 0.03 m at z = 11.1 1. NO chain (test01BNC040.418 (1. as described above. From the analysis of the experimental envelope curve (solid line).157 (1.7 m pole with chain damper. the distinction between x and y components is omitted.156) 1. NO chain (test01BNC081.L suggests limited amplitude effects. A cross comparison of the two tables and for equivalent x0.x located between 2. Caracoglia. Values of ζ1. deviations were linked to the non-linearity of the device.1% and 0. 5(b)) selected for the identification of A(t).x = 4.1%.25 2. Note 2: Measured f in parentheses computed during free-vibr. (a) In-plane displacement time-history at z = 11.142 n.24 m before release). the “whirling” of the chain may be more quickly initiated.1 1.0% depending on the different configurations and amplitudes.y without first-mode damper were derived from residual behavior in the presence of hanging chain.109 n.461) 1. substantially confirming the original estimates. 5. M L = 0.417 n. note 1) M L = 0.8 m.y derived from residual values in the presence of hanging chain. 5(b) (dashed thick line).). note 1) M L = 0.468) 1. canister damper removed.5% and 4.1 −0. This example corresponds to large oscillation amplitudes (0.8 m (digitally band-pass filtered). Table 3(a) and (b) summarize the results related to the computation of first-mode damping through Hilbert transform for different pole configurations without canister damper (tests conducted before the removal of the canister damper confirmed the findings reported in these tables). it was decided to conservatively evaluate the performance of the chain from a design perspective by considering an averaged value accounting for both regions of active (t < 13 s in Fig. x0. z = 11.L > 0. aluminum-alloy light poles. Washington (DC.National Research Council.107(7):1550–69. . 91(9):1163–73. Engineering Structures 2001. in which the inactive-chain behavior is accounted for in the calculation of the equivalent damping ratio.. New York (NY. Acknowledgements This research was supported by the Illinois Department of Transportation (Grant IHR-R27. Research Report. at the same time. technical supervisors in the NSEL and their staff was much appreciated during the design and development of the tests. 3-D gust effect factor for slender vertical structures. [4] Goswami I.41(1–3):357–68. The collaboration of Dr. Scanlan RH. Background information regarding the luminaire failures on the I-80 Le Claire Bridge over the Mississippi River and the pole failures in Galesburg and Woodhull Areas. N. thicker and shorter units) [23]. including collapse limit states governed by along-wind buffeting. Turbulence modeling for gust loading. Jones NP.17(2):143–55. New York (NY. Journal of the Italian Technical Association for Steel Construction (CTA) 1999. Jones / Engineering Structures 29 (2007) 821–831 in Fig. 2004. [12] Mabie HM. Lippert of the Bureau of Materials and Physical Research. Jones NP. Journal of Wind Engineering and Industrial Aerodynamics 1992. Standard specifications for structural supports for highway signs. USA): University of Illinois at UrbanaChampaign. Finally. Wind effects on structures. the investigation demonstrates that. Concluding remarks This paper presents the results of a recent analytical. 2003–2004) under the technical direction of Mr. [18] Repetto MP.(LI(1)):39–51 [in Italian]. although vulnerable to wind-induced failures associated with this particular event. Washington (DC. Probabilistic Engineering Mechanics 2002. despite the limits on structural damping at large amplitudes and the presence of three-dimensional motion effects partially neutralizing the effects of the proposed damper.P. luminaires and traffic signals. [5] Blevins RD. could be responsible for large vibration amplitudes (aeroelastically related or enhanced). Since the nature of the oscillations was clearly affected by insufficient structural damping in the fundamental mode. Minimum design loads for buildings and other structures (ASCE702). [2] IL-DOT . [15] Caracoglia L. Solari G. Solari G. 1990. USA. 1975. Transverse vibrations of double-tapered cantilever beams with end support and with end mass. Journal of Fluids and Structures 1999. New York (NY. Free-vibration analysis of beams and shafts. USA): Transportation Research Board . 2003. Journal of the Acoustical Society of America 1974. [6] Wieringa J. References [1] AASHTO. Updating the Davenport roughness classification. Roughness estimation for wind-load simulation experiments.55(5):986–9. 4th ed. [14] Gorman DJ. [3] Simiu E. Dynamic alongwind fatigue of slender vertical structures. USA): American Society of Civil Engineers. [13] Mabie HM. numerical and experimental investigation focused on the analysis of multiple collapses experienced by highway light poles during a wind storm in Western Illinois. [16] Pagnini LC. 10. can be employed as an alternative to steel units. Grzegorz Banas and Mr. Reston (VA. the proposed mitigation solution considered the installation of an impact damper (hanging chain). [19] Kaczinski MR. Experiments were performed at the University of Illinois. Newmark Structural Engineering Laboratory (NSEL). this study and supplemental investigations concluded that. 2002. or Federal Highway Administration. The study showed that testing is crucial in these situations. Journal of the Acoustical Society of America 1972. In addition. This conclusion was derived after carefully analyzing other possible causes of failure. All these mechanisms were excluded since they did not seem compatible with the particular meteorological condition and. Gust buffeting and aeroelastic behaviour of poles and monotubular towers. optimized for the first mode. 2nd ed. Documentation transmitted to UIUC on August 21. “Costruzioni Metalliche”.23(12):1622–33. Lagomarsino S. if appropriately designed and without the addition of the proposed first-mode damper depending on the geometry and configuration (e. Timothy J. Van Dien JP.g. 1998. It was concluded that. 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