Engineering Structures 34 (2012) 173–186

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Engineering Structures

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Field testing and analysis of aluminum highway sign trusses

Jennifer A. Rice a,⇑, Douglas A. Foutch b, James M. LaFave b, Schaun Valdovinos c

a Department of Civil and Coastal Engineering, Engineering School of Sustainable Infrastructure and Environment, University of Florida, 365 Weil Hall, P.O. Box 116580,Gainesville, FL 32611, United Statesb Department of Civil and Environmental Engineering, University of Illinois at Urbana-Champaign, 205 N. Mathews Avenue, Urbana, IL 61801, United Statesc Hatch Mott MacDonald Ltd., 1010-1066 West Hastings Street, Vancouver, BC, Canada V6E 3X2

a r t i c l e i n f o

Article history:Received 27 March 2011Revised 27 June 2011Accepted 7 September 2011Available online 4 November 2011

Keywords:Wind loadingSign support trussesFatigueAluminumDrag coefficient

0141-0296/$ - see front matter � 2011 Elsevier Ltd. Adoi:10.1016/j.engstruct.2011.09.021

⇑ Corresponding author. Tel.: +1 806 239 6838; faxE-mail addresses: jrice@ce.ufl.edu (J.A. Rice), dfout

jlafave@illinois.edu (J.M. LaFave), schaun.valdovinosnos).

a b s t r a c t

Four representative Illinois Department of Transportation (IDOT) highway sign support trusses wereevaluated analytically and in the field to assess their design load capacity and fatigue resistance whensubjected to wind loading. The three overhead span bridge and one cantilever welded aluminum trussstructures were instrumented for static and dynamic field testing with strain gages and accelerometersto collect structural behavior data when subjected to manual excitation, wind loading, and truck gustexcitation. Companion analytical models were developed and calibrated to represent the structures inthe field, after which these were then used to assess the response of the trusses at full design wind loads.Current IDOT and AASHTO highway sign support structure design loading and resistance approaches withrespect to both strength and fatigue have been critically assessed in light of the analyses and field testresults, with recommendations made for improving future designs.

� 2011 Elsevier Ltd. All rights reserved.

1. Introduction Research Program (NCHRP) funded two projects, the first related to

Aluminum cantilever and overhead sign structures (OSS) havelong been used to display signs and other signals requiring highvisibility from roadways. In the past, the state of Illinois Depart-ment of Transportation (IDOT) has had some difficulty with certainof their sign support truss structures consisting of welded hollowtube sections, which were susceptible to crack formation at someof the web-to-chord welded connections. In particular, IDOTobserved cracking at ‘‘T’’, ‘‘Y’’, and ‘‘K’’ tube-to-tube welded con-nections, most likely as a result of wind-induced cyclic loading.In the most extreme cases, complete connection failures resultedin web members falling on to the roadway [6]. In an effort to mit-igate wind-induced fatigue damage, IDOT changed their standarddesigns for aluminum highway sign structures in the year 2000.The design changes included eliminating redundant members,increasing the diameter and thickness of the hollow circular tubes,and installing vibration dampers, all with the intention of increas-ing the strength, stiffness, and factor of safety against failure forthese structures.

Wind damage to overhead and cantilever sign structures is notunique to Illinois; many states have reported damaged and/or col-lapsed sign structures [7,8,18]. The problem has been so broad forcantilevered sign structures that the National Cooperative Highway

ll rights reserved.

: +1 352 392 3394.ch@illinois.edu (D.A. Foutch),@hatchmott.com (S. Valdovi

their wind response [10] and the other about design for wind loads[5]. Results from the latter report were used in developing the fourthedition of the AASHTO Standard Specifications for Structural Supportsfor Highway Signs, Luminaries and Traffic Signals (2001) [1].

A series of full-scale measurements and analyses were con-ducted to assess the new IDOT sign support designs. The objectivesof this research were fourfold: (1) determination of the AASHTOdesign load capacity and fatigue resistance of representative IDOTsign structures based on analysis and testing, (2) development ofreliable analytical models for computing the response of aluminumhighway sign truss structures to natural wind and truck-inducedloads, in support of the first objective, (3) evaluation of the effec-tiveness of vibration dampers installed on the sign trusses, and(4) identification of reasons for observed cracking of certain mem-bers at connections in older IDOT sign trusses. Detailed informa-tion on laboratory and field testing of the dampers can be foundin Rice and Foutch [14], and an evaluation of older IDOT signtrusses (including features that made them susceptible to fatiguecracking) is given in Rice et al. [15]. This article summarizes theexperimental approach and results from the study, with a specificemphasis on full-scale field testing, analytical modeling, andassessment of current design guidelines; additional details maybe found in an IDOT final project report [6].

2. IDOT sign supports

The current IDOT [9] design guide identifies four basic signstructure types – the cantilever structure (II-C-A) and Types I-A,

174 J.A. Rice et al. / Engineering Structures 34 (2012) 173–186

II-A, and III-A overhead sign bridges. The particular sign bridge des-ignations (and therefore their aluminum member sizes) are basedon the space truss span and the area of the sign that it supports.The IDOT overhead support structure selection criteria are summa-rized in Table 1.

One of each of these four types of structures was studied in thisproject; the cantilever and Type III-A structures included in thestudy are shown in Fig. 1. The analytical and experimental studiescarried out for each sign support structure type were fairly similar.Therefore, the overall approach and results are mainly presentedhere for the Type III-A OSS (the largest one tested), as a represen-tative example.

IDOT OSS are comprised of four-chord aluminum trusses sup-ported by vertical planar steel column trusses. The main trussesare made up of multiple truss sections (as shown in Fig. 2) con-nected by bolted flange-splice connections. The truss designatedby the IDOT Bureau of Bridges and Structures as a Type III-A isthe largest truss of the overhead type, measuring 1.52 m (5 ft) wideby 2.13 m (7 ft) tall (center-to-center of the chords).

The geometry and properties of the four sign support trussesinvestigated in this study are summarized in Table 2. All trusschords and web members have wall thicknesses of 7.93 mm (5/16 in.). The total spans given are from center-to-center of supports.

All of the support trusses evaluated have total sign areas thatare smaller than that permitted for their type and span. Becausewind acting on the sign faces is by far the most significant contri-bution to loading of these structures, it would therefore be ex-pected that measured stresses in the tested support trusses willbe lower than allowable design stresses (even at the design windloading) due to the relatively smaller sign areas. However, the ac-tual measured responses were still significant enough for all thepurposes of this study (analytical model validation, peak designstress estimation, drag coefficient and gust factor assessment,damping estimation, fatigue assessment, etc.).

3. Design loads

3.1. Wind loads

Design wind loads for sign structures are calculated accordingto the Standard Specifications for Structural Supports for HighwaySigns, Luminaires and Traffic Signals [1]. Wind loading is consideredas a static pressure acting horizontally on a sign and on the rest ofthe structure, calculated in accordance with that specificationusing a design wind loading expression, resulting in psf (where1.0 psf = 0.048 kPa), as follows:

P ¼ 0:00256KzGV2ICdðpsfÞ ð1Þ

where P = design pressure, Kz = velocity pressure exposure coeffi-cient, G = gust effect factor, V = basic wind speed (mph), I = impor-tance factor, Cd = drag coefficient.

The velocity pressure coefficient, Kz, is taken as 1.0 for WindExposure Category C (open terrain with scattered structures) anda maximum sign mounting height of 10 m (32.8 ft), while theimportance factor, I, is taken as 1.0 (for a 50-year design life) forall of the sign truss structures.

Table 1IDOT overhead sign truss selection criteria.

Overhead signtruss type

Maximum span lengthrange [m (ft)]

Maximum sign area range[m2 (ft2)]

I-A 21.4–30.5 (70–100) 32.6–56.7 (350–610)II-A 27.5–39.7 (90–130) 68.8 (740)III-A 36.6–48.8 (120–160) 83.7–111.6 (900–1200)

The basic wind speed, V, is defined as the three-second gustwind speed at a height of 10 m (32.8 ft) above ground. This takeson regional values, corresponding to a mean return period of50 years – for Illinois, the basic wind speed is 40.2 m/s (90 mph).

The gust effect factor, G, used in Eq. (1) accounts for two phe-nomena related to the dynamic interaction between the structureand the gustiness of the wind. One is the anticipated spatial varia-tion of the wind pressure acting on a structure, and the other is thegustiness of the wind with respect to the three-second basic windspeed. The response that is calculated when using the design codewind is the static response to the three-second average windspeed, increased to account for gust magnitudes and dynamicamplification through the use of the gust effect factor. In a flexibleand relatively small structure, such as a highway sign truss, spatialvariation of the wind pressure is of little effect, whereas dynamiceffects will be shown to be significant, especially for low to moder-ate wind speeds. The gust effect factor therefore adjusts the effec-tive wind pressure to account for this interaction between thestructure and the wind gusts in an average sense. Calculation ofG for dynamically sensitive structures can be quite a complicatedprocess [3], so design specifications have traditionally adopted avalue, based on historical research results, which has performedwell. This value (G = 1.14) is used in conjunction with the basicwind speed (three-second gust) to obtain an effective velocitypressure for use in design calculations [1]. The gust effect factorwill be further examined later on in this article.

Pressures acting on the surface of a structural body are the re-sult of the wind velocity directly outside the boundary layer. Forthe same wind velocity, different bodies will experience differentpressures due to geometry and their Reynolds numbers. The intro-duction of a non-dimensional drag coefficient, Cd, accounts forsuch differences in applied force. Because of the difficulty in quan-tifying the wind pressures that actually result from turbulent windflow across a variety of structures and components, a lot of exper-imental data should be generated and utilized to adequately deter-mine drag coefficients acting on various types of structures. Theseexperiments are typically conducted via full-scale testing, model-ing, and wind tunnel tests [17]. The first work done to accuratelydetermine drag coefficients on wall panels mounted at the groundwas conducted in the 1930s by Flachsbart. Drag or force coeffi-cients were determined as a function of the aspect ratio (width/height) of the panels. This research became the basis of many de-sign codes, but it did not consider turbulent flow or the effects ofpanels raised above ground level (as is the case for highway signs)[17].

When a sign panel is located a significant distance up from theground, the wake flow is dominated by separating horizontal shearlayers above and below the sign. This separating shear layer inter-action increases the drag coefficient. Therefore, the appropriatedrag coefficient to be used for a solid rectangular sign has beenexamined by Letchford [12]. Results of wind tunnel tests show thatthe drag or force coefficient is not only a function of the aspect ra-tio, B/s, but also the ‘‘clearance ratio,’’ s/h (see Fig. 3). The ASCE 7-05(2006) general loading design standard has been conservativelyupdated to reflect this phenomenon.

Drag coefficients from the AASSHTO [1] specification, the IDOT[9] design manual, Letchford’s [12] tests, and ASCE 7-05 [3] are allprovided in Table 3.

Allowable stresses for this type of sign structure can be deter-mined according to the AASHTO design specifications (2001). Thecode increases the allowable stress by 33% for certain load combi-nations including wind; in tension the adjusted allowable stress is101 MPa (14.63 ksi), and in compression it is 110 MPa (15.96 ksi).These allowable stresses are still considerably lower than the min-imum yield strength of aluminum [241 MPa (35 ksi)], due both tobuilt-in safety factors and the fact that the effective yield strength

Fig. 1. IDOT sign support trusses: (a) cantilever-type and (b) overhead Type III-A.

Fig. 2. Isometric view of a typical truss section.

Table 2Properties of the trusses included in the experimental study.

Truss type Total span[m (ft)]

Number of truss units Total sign area[m2 (ft2)]

Percentage of max. allowedsign area

Truss chords Web membersOutside diameter[mm (in.)]

Outside diameter[mm (in.)]

II-C-A 9.1 (30) 1 14.2 (153) 45 165 (6-1/2) 82.6 (3-1/4)I-A 26.8 (88) 3 32.7 (352) 58 127 (5-0) 63.5 (2-1/2)II-A 35.0 (115) 3 47.7 (513) 69 178 (7-0) 76.2 (3-0)III-A 42.3 (142) 4 52.0 (560) 50 216 (8-1/2) 88.9 (3-1/2)

J.A. Rice et al. / Engineering Structures 34 (2012) 173–186 175

of the material within an inch of connections is further reduceddue to welding [11].

Design loads prescribed by the IDOT [9] specification have beenderived from a 40.2 m/s (90 mph) basic wind speed, a gust effectfactor of 1.14, a drag coefficient of 1.19 (rounded to 1.2), and animportance factor of unity. A conservative simplification in the IDOTdesign manual prescribes that the wind loading be applied to astructure as shown in Fig. 4, with a horizontal pressure applied tothe sign panel area of 1.4 kPa (30 psf) and with truss member

loading simplified by considering the exposed truss as a closed re-gion with an average applied pressure of 0.48 kPa (10 psf). However,to enable the most accurate comparison here with measured (orextrapolated) values from the field tests, the actual loads resultingfrom the appropriate drag coefficients and wind loads have beencalculated and applied to the signs and trusses, resulting in distrib-uted loads along the exposed tubular truss members. Those result-ing loads are summarized in Table 4 (for the case when windvelocity is taken equal to the basic wind speed of 90 mph).

Fig. 3. Sign panel clearance ratio and length labels.

176 J.A. Rice et al. / Engineering Structures 34 (2012) 173–186

3.2. Fatigue loads

When a structure or component is subjected to repeated stressvariations, even at stress levels well below the allowable stress ofthe material, small cracks may form. With continued cyclic loading,these cracks may propagate, leading to failure of the component orstructure. Such fatigue cracks tend to form in regions of stressconcentrations, such as at notches, holes, welds, or other discontinuities.

The AASHTO [1] design specifications for fatigue are based inpart on NCHRP Report 412, Fatigue Resistant Design of CantileveredSignal, Sign and Light Supports [10]. When following that ap-proach, components of a sign structure should be designed forfatigue to resist appropriate equivalent static loading due to gal-loping, natural wind gusts, and truck-induced gusts. Stressesmust not be greater than the Constant Amplitude Fatigue Limit(CAFL) listed by AASHTO for each detail category. The CAFL val-ues applicable to members tested in these sign structures inthe field are 13.1 MPa (1.9 ksi) for chord members and 3.0 MPa(0.44 ksi) for web members. AASHTO specifications require fati-gue design of cantilever sign structures, but not necessarily forspan-type structures (though they have now been included inthe latest version).

Two loading conditions are considered for AASHTO sign trussfatigue design: natural wind gusts and gusts induced by truckspassing under the structure. In both cases an equivalent static windpressure is calculated and applied to the structure, and the result-ing member stresses are then compared to the appropriate CAFLvalue. While such a simplified fatigue analysis approach is cur-rently considered acceptable for design from the standpoint ofAASHTO, it may in fact not necessarily always be conservative orable to provide an accurate assessment of fatigue vulnerability.Emerging rational approaches have been developed that moveaway from simply using reference fatigue design wind pressuresin conjunction with the CAFL, and toward a more robust account-ing of the cumulative effects of wind occurring at various velocities[16,13].

Per AASHTO, a sign truss is designed with respect to fatigue toresist the following equivalent static natural wind pressure, PNW:

PNW ¼ 5:2CdI ðpsfÞ ð2Þ

Table 3Recommended drag coefficients as a function of aspect and clearance ratios.

Sign truss Clearance ratio (s/h) Aspect ratio (B/s) Cd

AASH

Cantilever (II-C-A) 0.36 1.04 1.12Type I-A 0.49 1.59 1.16Type II-A 0.50 2.53 1.19Type III-A 0.38 3.47 1.19Single sign 0.30–0.50 1.0–1.5 1.12–

This formula is based on Eq. (1), using a yearly mean wind speedof 5 m/s (11.2 mph) to represent the fatigue-limit-state windvelocity, which is defined as having a 0.01% probability of exceed-ance assuming a Rayleigh distribution [10].

Similarly, the equivalent static truck gust pressure, PTG, for fati-gue design is given by:

PTG ¼ 18:8CdI ðpsfÞ ð3Þ

This equivalent truck gust pressure is to be applied in the verti-cal direction to the truss, as well as to all portions of the structureprojected onto a horizontal plane. Eq. (3) is the result of the windpressure formula given in Eq. (1) with a wind speed of 30 m/s(65 mph), set to coincide with typical posted traffic speed limits[1]. This formula currently used by AASHTO for static truck gustpressure in fatigue design was reduced from an earlier version(PTG = 36.6CdIf, in the 2001 AASHTO specifications) starting withthe 2002 AASHTO interim revisions [2].

4. Analytical models

Finite element models were created using SAP2000 [4] for eachof the evaluated sign structures based on design and shop draw-ings, as well as project construction specifications, provided byIDOT (dimensions were also later field-verified). Members wererepresented by beam elements that were assigned appropriatedimensions and material properties, and all truss chord-to-chordand web-to-chord connections were assumed to be rigidly framed.

Initial models were created prior to instrumentation and testingof the in-service sign trusses to assess their expected response anddetermine optimal sensor placement. Following testing and analy-sis of the measured dynamic response (as presented in a subse-quent section), the models were updated to even better reflectthe behavior of the real structures. Most of the uncertainty in themodels was with respect to their stiffness, attributable to acciden-tal eccentricities due to fabrication imperfections, conditions ateach end of a truss where it is connected to the supporting struc-ture (and at the ends of all truss web members), and end fixity atthe bottom of the support frames. For instance, the ends of a trussconnect to column support structures via four U-bolts, so friction isthe primary resistance to longitudinal movement of a truss relativeto the supports (I-beams of the support structure are used for seat-ing the truss chords, which bear directly on their top flanges). Thefirst attempt to model each structure usually resulted in a modelthat was a bit too stiff, as indicated by the calculated natural fre-quencies being higher than the measured ones.

The static and dynamic finite element analysis results presentedherein incorporate the updates made after initial field testing andare also used to evaluate the anticipated stresses resulting fromthe statically applied design loads (according to AASHTO specifica-tions). Complete details of the modeling process, includingassumptions regarding the truss-to-support connections, may befound in the IDOT final project report [6]. In general, slightlychanging the support conditions in the models had some affect

TO (2001) IDOT (2007) Letchford (2001) ASCE 7-05 (2006)

1.20 1.39 1.781.20 1.40 1.721.20 1.43 1.701.20 1.47 1.76

1.16 1.20 1.39–1.44 1.75–1.80

Fig. 4. Design wind loading [9].

Table 5Modal analysis results – Type III-A truss.

Mode Modedescription

Model frequency(Hz)

Measured frequency(Hz)

1 Longitudinal 1.12 1.232 Transverse 1.76 1.853 Vertical 2.61 2.554 Torsional 4.52 5.09

J.A. Rice et al. / Engineering Structures 34 (2012) 173–186 177

on the calculated natural frequencies, but little impact on the cal-culated stresses.

4.1. Dynamic behavior

Results of the finite element modal analysis for the Type III-Asign structure are given in Table 5, with each mode shape de-scribed below and also illustrated in Fig. 5. Field testing to deter-mine the dynamic properties of each sign are discussed in moredetail later; however, those measured results are also presentedin Table 5, to show how closely the model represents the dynamicbehavior of the in-service structure. For this structure, all the mod-eled natural frequencies except for the vertical mode are slightlylower than the measured frequencies; however, based on observa-tions during the model updating process, the differences have anegligible impact on the resulting member stresses. These modeshapes and frequency ranges are typical for all of the OSS trussesevaluated in this study.

Mode 1: longitudinal motion is developed almost entirelythrough deformation of the support structures. This mode is notexcited to any considerable extent by wind gusts, due to the slimtruss profile in this direction.

Mode 2: the most important mode for the sign structure istransverse horizontal motion, which is easily excited by windand truck gust loading on the face of the signs (and can be seenin the plan view below).

Mode 3: vertical motion is a combination of bending in the trussand the columns. Both wind and truck gusts can excite this mode.

Mode 4: there is considerable torsion about the longitudinalaxis of the truss under wind loading. This mode has a higher fre-quency relative to the first three fundamental modes.

4.2. Static loading response

The stresses shown in Table 6 are results from the finite ele-ment model using the ‘‘design’’ (actual applied) wind loading(given in Table 6) and dead load acting together on the Type III-Asign truss. As expected, axial forces dominate due to truss action.Chord members experience the highest stresses; however, theyare much smaller than the yield stress of 241 MPa (35 ksi), and

Table 4Drag coefficients and resulting member wind loads for evaluated sign trusses.

Element II-C-A I-A

Cd Load Cd Load

Chord 0.82 135.1 N/m (9.26 lb/ft) 1.10 138.6 N/m (9.50 lb/Diagonals 1.10 90.2 N/m (6.18 lb/ft) 1.10 69.3 N/m (4.75 lb/ftSign 1.20 1.19 kPa (24.9 psf) 1.20 1.19 kPa (24.9 psf)

even well below the allowable stresses. Horizontal diagonal ele-ments also experience high stresses relative to other members,as they are responsible for transferring transverse forces from thechords on the front (sign) face of the truss to the chords on the backface of the truss.

Table 7 illustrates the effect of using the higher drag coefficientssuggested by Letchford [12] and ASCE 7-05 [3]; peak stresses in-crease by up to 22% and 45%, respectively.

Similar results of modeled static loading response were foundfor the other evaluated OSS trusses, which indicates that the webmembers most susceptible to overload (and possibly also to dam-age from cyclic loads) are the horizontal diagonals (HD). These re-sults also informed the instrumentation plan described in thefollowing section.

5. Field testing

The purpose of the field investigation was to measure the actualresponse of sign structures under various types and levels of exci-tation. The primary sources of loading on the structures are wind(including gusts) acting on the sign panels and short bursts of exci-tation from trucks and other vehicles passing below them. Thestructures were instrumented to capture acceleration and strainunder these loading conditions, as well as during manual excita-tion tests. An anemometer was placed at the site of each structureto quantify the wind speed and direction, for correlation with re-sponse levels.

5.1. Instrumentation

To ensure that the response of each structure was captured dur-ing times when the average wind velocity normal to the sign panelexceeded approximately 13 m/s (30 mph), sensors were left on thestructures over a period of several months. Ends of the sensorcables were bundled in a protected box mounted at each structure,and when it was time to acquire data, the cable ends were accessedand connected to data acquisition hardware and a laptop computerthat were transported to the site at the time of data collection. Thisapproach enabled timely acquisition of pertinent structural re-sponse without the need for lane closure or other disruption totraffic.

To facilitate simultaneous instrumentation of two of the struc-ture types, two sets of instrumentation were used during fieldtesting. Disposable foil strain gages were also used in all cases. TerraTechnology accelerometers with a sensitivity of 2.5 V/g and a rangeof ±5 g were utilized on the Type II-C-A and Type II-A sign

II-A III-A

Cd Load Cd Load

ft) 0.75 132.2 N/m (9.06 lb/ft) 0.58 124.8 N/m (8.55 lb/ft)) 1.10 83.2 N/m (5.70 lb/ft) 1.10 97.0 N/m (6.65 lb/ft)

1.20 1.19 kPa (24.9 psf) 1.20 1.19 kPa (24.9 psf)

Fig. 5. (a) First mode, elevation view of longitudinal motion, (b) second mode, plan view of transverse horizontal motion, (c) third mode, elevation view of vertical motion,and (d) fourth mode, profile view of torisonal motion.

Table 6Calculated member stress in the Type III-A structure for dead load plus applied wind load (excluding G).

Member Axial stress [MPa (ksi)] Bending stress [MPa (ksi)] Max. Total stress [MPa (ksi)] % Axial

Chord 1 44.3 (6.43) 4.2 (0.61) 48.5 (7.03) 91Chord 2 27.1 (3.93) 5.2 (0.76) 32.3 (4.69) 84H1 4.68 (0.68) 0.62 (0.09) 5.30 (0.77) 88HD 22.3 (3.23) 2.00 (0.29) 24.3 (3.52) 92V1 13.8 (2.01) 1.2 (0.17) 15.0 (2.18) 92V2 16.4 (2.38) 1.86 (0.27) 18.3 (2.65) 90

178 J.A. Rice et al. / Engineering Structures 34 (2012) 173–186

structures, while PCB accelerometers with a sensitivity of 1 V/g anda range of ±5 g were used on the Type I-A and Type III-A trusses. AWindMaster ultrasonic anemometer (Gill Instruments), capable ofmeasuring three axes of wind speed from 0 to 70 m/s(0–157 mph) was installed on the Type II-C-A and Type II-A signstructures. The anemometer installed on the Type I-A and TypeIII-A trusses was a WindMonitor propeller-type unit (Young), mea-suring horizontal wind speed [up to 44.7 m/s (100 mph)] and direc-tion (0�–360�). To ensure that ambient wind characteristics werecaptured at the site, with as little interference from the structureas possible, the anemometers were installed at least 2 m abovethe top chords of the truss, over one of the support columns.

Table 7Comparison of calculated member stress in the Type III-A structure for dead load plus ap

Member IDOT Cd = 1.20 [MPa (ksi)] Letchford Cd

[MPa (ksi)]

Chord 1 48.5 (7.03) 57.2 (8.30)Chord 2 32.3 (4.69) 37.6 (5.45)H1 5.30 (0.77) 6.31 (0.92)HD 24.3 (3.52) 29.6 (4.29)V1 15.0 (2.18) 16.0 (2.33)V2 18.3 (2.65) 19.6 (2.84)

The data acquisition system consisted of a National Instrumentssignal conditioning chassis capable of simultaneously capturingstrain (in a quarter Wheatstone bridge configuration), acceleration(with an 8-pole elliptical anti-aliasing filter), and wind speed/direction. A sampling rate of 100 Hz was used, resulting in a mea-surement bandwidth of 50 Hz. LabView software was used for datacollection, visualization, and storage.

Placement of strain gages was selected based on regions ex-pected to experience higher stress, per results from the finite ele-ment analyses. In the case of the overhead trusses, these regionswere on truss web members closest to the supports and on maintruss chords at mid-span. Although the locations of the signs

plied wind load for various recommended drag coefficients.

= 1.47 ASCE 7-05 Cd = 1.76

% Incr. [MPa (ksi)] % Incr.

18.1 66.7 (9.67) 37.516.3 43.3 (6.27) 33.720.4 7.46 (1.08) 42.421.8 35.2 (5.11) 45.26.71 17.1 (2.48) 13.97.05 20.9 (3.04) 14.6

J.A. Rice et al. / Engineering Structures 34 (2012) 173–186 179

mounted on the overhead trusses are not perfectly symmetric,strain gages on members near the supports were only placed atone end of each truss; measured stresses were then always com-pared to computed stresses from the same end of the model. Wher-ever site conditions allowed, strain gages were attached on thetruss end that was expected to experience higher stresses (i.e., onthe end with the higher percentage of sign area), such as at theright end of the Type III-A truss shown in Fig. 1b. For the cantilevertruss, the key high-stress regions were truss web members andmain truss chords adjacent to the support column. At each selectedstrain gage location, two gages were placed, one on either side ofthe tubular section, so that both the horizontal bending and axialcomponents of stress could be resolved. The focus was on horizon-tal (transverse) bending of the members since wind loading is theprimary load of interest. Table 8 summarizes the strain gagelocations.

Accelerometer placements were selected so the primarybending and torsional mode shapes of a truss could be captured.Accelerometers were only placed on half of the truss span, at theend-, quarter-, and mid-points (on both a top and bottom chord).The layout of horizontal (transverse), vertical, and longitudinal

Table 8Strain gage plan, Type III-A sign truss.

Strain gage Location

C1a Bottom chord, front planeC1bC2a Bottom chord, back planeC2bH1a Horizontal diagonal, top planeH1bHDa Horizontal diagonal, bottom planeHDbV1a Vertical diagonal, front planeV1bV2a Vertical diagonal, back planeV2b

Fig. 6. Accelerometer placement

accelerometers for the Type III-A structure, which is representativeof the others, is summarized in Fig. 6. Figs. 7 and 8 show typicalrepresentative accelerometer, strain gage, and anemometer instal-lations on truss elements. The anemometer is mounted so thatwhen the wind direction (and thus the anemometer alignment)is parallel to the truss chords, the wind direction reading is either0� or 180�; wind acting normal to the back of the sign results in a90-degree wind direction reading, as indicated in Fig. 8.

5.2. Dynamic response

Manual excitation tests were performed on the trusses for thepurpose of verifying modal analysis results from finite elementmodeling. Tests were conducted for both horizontal and verticalvibration of each truss. In the horizontal direction, the truss wasexcited by a person standing at the center of the truss span andmoving their body back and forth until a substantial vibrationamplitude was reached. Vertical tests were performed using a ropeattached at mid-span of the truss, which was pulled down repeat-edly by a person standing below the truss until a high amplitude ofvibration was reached. In both cases, the truss was then allowed tovibrate freely as the motion dissipated. Measured natural frequen-cies from the manual excitation tests and from subsequent windexcitation tests on the Type III-A structure were given in Table 5– they match well with those resulting from the final, updated, fi-nite element model of the structure, which was typically the caseacross all four sign truss structures tested in the field.

Each of the sign trusses has a Stockbridge-type damper installedfor the purpose of mitigating excessive vibration. One objective ofthis study was to assess the effectiveness of the dampers and makerecommendations for their improved usage. So, in addition toestablishing the fundamental frequencies of truss vibration, man-ual excitation tests also enabled comparison of structural dampingwith the dampers engaged vs. disengaged, in both the transverseand vertical modes. In general, the dampers added somewhat tooverall structural damping (in all but the Type III-A truss), asshown in Table 9 for the transverse horizontal mode damping.

for the Type III-A sign truss.

Fig. 7. Sensors mounted on a sign truss – accelerometers (left) and a strain gage (right).

Fig. 8. Aerial view of the Type III-A sign truss showing the alignment of the anemometer, and a picture of the anemometer mounted above the east support of the truss (topright).

Table 9Effect of Stockbridge dampers for the transverse mode on each sign truss type.

Truss type Percent of critical damping Percentage increase

Without damper With damper

II-C-A 1.21 1.99 64I-A 0.80 0.84 5II-A 1.42 1.61 12III-A 0.80 0.80 0

180 J.A. Rice et al. / Engineering Structures 34 (2012) 173–186

Additional laboratory tests were conducted to determine how thedampers could be modified to improve their performance bymatching their natural frequencies more closely to the fundamen-tal frequencies of the sign structures. More details about results ofthese tests and related laboratory experiments can be found in Riceand Foutch [14].

Response of a sign support structure to wind gusts is oftenthought of as being nearly static in nature, with the dynamic por-tion very small relative to quasi-static response. A plot of axialstress in a chord member of the cantilever structure, along withthe three-second moving average of axial stress, is shown inFig. 9 (recall that a three-second average wind speed was used

for calculating design loads), with an expanded region given toshow a detail of the two curves. The figure clearly demonstratesthat the dynamic component of response is not negligible. It shouldfurther be expected that the ratio of peak dynamic response di-vided by the three-second average would tend to be quite smallas the wind velocity gets large, but would become fairly large asthe three-second average approaches zero.

To assess the variation in average peak response as a function ofthree-second average response at various wind speeds, the ratio ofaverage peak response to three-second average response was cal-culated for a large number of data points from all the sign struc-tures. The peak-to-average ratios at five wind velocities are givenin Table 10.

All this data was combined into a single relationship for assess-ing the design of the structures. The resulting curve of response ra-tio vs. wind speed is shown in Fig. 10 (left). If the ratios areextrapolated up to the 40.2 m/s (90 mph) design wind speed, theaverage ratio would be about 1.05. While it is difficult to draw a di-rect comparison between the response ratio and the gust effect fac-tor, G, this value is on the same order as the G value used in theAASHTO specifications. Also shown in Fig. 10 is that the averageextrapolated ratio at the wind speed used for fatigue design

Fig. 9. Axial stress in cantilever chord member, and three-second average, vs. time (left), and detail of the boxed region (right).

Table 10Average ratio of peak response to three-second average (vs. average wind speed).

Sign Truss Wind Speed [m/s (mph)]

4.5 (10) 6.7 (15) 8.9 (20) 11.2 (25) 13.4 (30)

II-C-A – 2.63 1.82 1.66 1.39I-A – 2.39 1.63 1.40 1.33II-A – 2.61 1.80 1.38 –III-A 2.10 1.95 1.77 1.52 1.38

Table 11Statistical properties of response ratios.

Wind speed [m/s(mph)]

Mean StandardDeviation

Coefficient ofVariation

l r cv = r/l

6.7 (15) 2.27 0.784 0.3458.9 (20) 1.76 0.526 0.29811.2 (25) 1.47 0.251 0.17113.4 (30) 1.38 0.146 0.106

J.A. Rice et al. / Engineering Structures 34 (2012) 173–186 181

[5 m/s (11.2 mph)] is 3.1, which certainly could be of significancewhen considering fatigue. A curve showing one standard deviationbounds (based on the experimental results) is then given in Fig. 10(right), with statistical properties of the response ratio data givenin Table 11; note that the variation becomes small for higher windspeeds, but large for lower wind speeds. As a result, use of an aver-age value of gust effect factor for strength design is probably quiteacceptable; however, using the average value for fatigue evalua-tions may require some further consideration. The larger variationin response ratio at lower wind speeds could also in part be attrib-uted to the higher signal-to-noise ratio at lower measured strains.

Two separate formulations for G are given in ASCE 7-05 – onefor rigid structures (fundamental frequency >1 Hz) and one forflexible structures (fundamental frequency <1 Hz). While theAASHTO specification assumes that all structures to which it

Fig. 10. Combined plot of mean response ratio for all sign structures (lef

applies are considered flexible, all of the sign support structuresin this study actually had natural frequencies above 1 Hz. In thespecific case of the Type III-A structure, G can be calculated as0.895 and 0.893 for the flexible and rigid formulations, respec-tively. The fact that these are less than one implies that this struc-ture is not expected to experience resonance and resulting loadamplification due to the lower frequency energy contained in thewind spectra, at least according to ASCE 7-05. It is important tonote, however, that ASCE 7-05 was primarily developed for build-ings. Sign truss structures are still quite flexible relative to buildingstructures, and thus significant vibrations at the natural frequencyof these structures were still observed (as seen in Fig. 9) that maywarrant using some type of amplification, or gust effect factor, indesign, such as the one prescribed in the AASHTO specification.

t), and the same plot with one standard deviation indicated (right).

182 J.A. Rice et al. / Engineering Structures 34 (2012) 173–186

5.3. Wind load response

To determine the Type III-A sign truss response to strong windevents, the acceleration, strain, and wind data were all collected ona day when the wind speed acting normal to the truss exceeded13.4 m/s (30 mph). Almost an hour of data was collected, withthe recorded normal component of the wind speed reaching ashigh as 17.1 m/s (38.2 mph). The wind speed and direction recordsfor that time period are presented in Fig. 11, with a wind directionof 90� corresponding to action normal to the back of sign face.

The turbulence intensity (TI) of the measured wind data wasdetermined by analyzing five-minute intervals of wind speed datain which the mean remained fairly constant. For each record, TIwas calculated by: TI ¼ ru=lu, where ru is the standard deviationof the wind velocity and lu is the mean of the wind velocity overthe same period of time. The average measured TI is 15.3%; themaximum and minimum are 13.2% and 17.7%, respectively, andthe standard deviation of the measured TIs is 1.8%. The value of tur-bulence intensity is not estimated or utilized in the AASHTO spec-ifications; however, it is determined per ASCE 7-05 and used in theequation for calculation of the gust effect factor, G, as described inthe previous section. According to ASCE 7-05, the turbulence inten-sity for Exposure Category C is given by:

TI ¼ I�z ¼ c33�z

� �1=6

ð4Þ

where c is the turbulence intensity factor (given as 0.2 for ExposureCategory C) and �z is the equivalent height of the structure (in ft). Forthe Type III-A sign structure, the equivalent height is the height tothe middle of the sign panel, or 8.3 m (27.2 ft), resulting in a TI of20.7%. These results indicate that the turbulence at this site wasfairly typical (i.e., close to the 0.2 turbulence intensity factor for thisexposure) during the testing period.

All of the strain records measured during truss field testing re-quired some assessment and adjustment prior to further analysis.One of the challenges associated with acquiring strain data on anin-service structure is the difficulty in identifying a point of zero

Fig. 11. Strong wind recorded

wind-induced strain. The strain gages are ‘‘zeroed’’ out at thebeginning of testing, but due to continuous wind and resultingstructural vibration, this does not necessarily represent a true stateof zero wind stress. The result is that the strain record contains anunknown offset from actual zero. Since the wind force acting on astructure, and therefore the resulting strain, is proportional to thesquare of the wind speed, a relationship between strain and windspeed could be developed by using regression analysis. From thisrelationship, an offset value (from zero) for each record was thendetermined. Fig. 12 shows curves of the three-second average windstress in each chord member as a function of normal wind speed.These curves were generated by fitting a quadratic polynomial tothe data such that the curve goes to zero when the wind velocityis zero. The boxed-in numerical values indicate the expected aver-age wind stress value for the design wind speed of 40.2 m/s(90 mph), values that are utilized in subsequent analysis and forcomparison with code design loading. Similar curves were alsodeveloped for the web members.

Extrapolated measured wind stresses and those calculated fromthe finite element model for several truss members are comparedin Table 12 (the model results exclude any use of the gust effectfactor, G, in the wind loading). Comparisons are made for eachmember at the actual location of the strain gages, which in somecases is not the exact location where the maximum strain/stressoccurred. The ‘‘measured’’ values compare quite well with thosepredicted by the model. Results from both the model and extrapo-lated measured stresses show that Chord 1 (on the front plane ofthe truss, where the sign is mounted), experiences higher stresslevels than Chord 2. This discrepancy between the front and backchord, which is more pronounced in the extrapolated measuredstresses, may be due to incomplete load sharing between the frontand back panels of the truss. Furthermore, the extrapolated fieldresults indicate higher stresses in Chord 1 than does the model; re-call that the loading applied to the sign in the model is calculatedusing a drag coefficient of 1.20, which is lower than the 1.47 valueexperimentally determined by Letchford [12] and the value of 1.76prescribed in ASCE 7-05 [13].

for the Type III-A truss.

Fig. 12. Type III-A wind stress in chord members vs. three-second average normal wind speed.

Table 12Type III-A comparison of extrapolated [to 40.2 m/s (90 mph)] measured wind stresses vs. model stresses.

Member Model stress – wind load only Extrapolated from measured stress

Axial stress [MPa(ksi)]

Bending stress [MPa(ksi)]

Max. total stress [MPa(ksi)]

Axial stress [MPa(ksi)]

Bending stress [MPa(ksi)]

Max. total stress [MPa(ksi)]

Chord 1 27.6 (4.00) 4.00 (0.58) 31.6 (4.58) 35.7 (5.17) 3.31 (0.48) 39.0 (5.65)Chord 2 22.5 (3.26) 3.93 (0.57) 26.4 (3.83) 18.8 (2.72) 4.60 (0.67) 23.4 (3.39)HD 23.7 (3.44) 5.79 (0.84) 29.5 (4.28) 21.6 (3.13) 1.93 (0.28) 23.5 (3.41)V1 8.21 (1.19) 2.62 (0.38) 10.8 (1.57) 3.52 (0.51) 1.03 (0.15) 4.48 (0.65)V2 3.17 (0.46) 1.45 (0.21) 4.62 (0.67) 3.86 (0.56) 1.86 (0.27) 5.72 (0.83)

Table 13Type III-A stresses due to measured (extrapo-lated) wind load + model dead load.

Member Total stress [MPa (ksi)]

Chord 1 38.6 (5.60)Chord 2 35.2 (5.10)HD 37.6 (5.46)V1 28.5 (4.14)V2 23.9 (3.47)

J.A. Rice et al. / Engineering Structures 34 (2012) 173–186 183

To verify the design calculations for wind loading, results fromthe analytical model plus field measurements were examined andcompared to allowable stress values. Total member stress, calcu-lated as the dead load from the model plus the values extrapolated[to the expected average peak stress at 40.2 m/s (90 mph)] frommeasured stresses due to wind are presented in Table 13. Thesevalues are all well below the lowest allowable stress value of101 MPa (14.63 ksi), for tension. One reason the stresses are lowcompared to the allowable design values is that the demand onthis OSS is lower than the demand it was designed for. The totalsign area currently installed on the truss is only about half of themaximum allowed, and the length of the truss is 2.4 m (8 ft) lessthan the allowed 45.7-m (150-ft) span; structures with sign areas

closer to the maximum allowed are expected to generate stressescloser to the design values. As an approximation of the stressesthat may be expected if the maximum allowed sign area and span

184 J.A. Rice et al. / Engineering Structures 34 (2012) 173–186

length are used, the chord and horizontal diagonal (HD) analyticalmodel and field-extrapolated stresses have been scaled by the ra-tios of the maximum span and sign area to the existing span andsign area, respectively; these results are shown in Table 14. Thescaled extrapolated chord stresses in all but the Type III-A truss ex-ceed the allowable tensile stress, indicating that there could besome cases of ‘‘overstress’’ if the maximum span and sign areaare utilized. Note that in all cases the extrapolated measured stres-ses exceed the model stresses, which is likely the result of the lowdrag coefficient used to calculate the sign pressure in the model.

5.4. Truck gust response

For each instrumented sign truss, the response (accelerationand strain) due to truck gusts was measured to evaluate the result-ing fatigue loads. As an example, for the Type III-A truss, which hada typical standard sign clearance above the roadway of 5.3 m(17.3 ft), truck gust measurements were obtained during portionsof two separate days when the ambient wind conditions were calm[less than about 4.5 m/s (10 mph)]. Over the course of the twodays, around 200 trucks were recorded passing under the truss.During these field tests, the type of truck was noted in order todetermine which types, if any, induced the greatest structural re-sponse. The type of trucks that were recorded include: typicaltrucks with single trailers, typical trucks with flow deflectors onthe cab, trucks with shorter boxy trailers, flatbed trailers, and tank-ers. There was no conclusive data as to which truck type causes thehighest response in this truss. The response is not only dependenton the truck type, but also on its travel speed, whether multiplevehicles are passing underneath, whether many vehicles in a roware passing underneath, and what the wind conditions are at thatparticular time.

Horizontal motion of the truss dominates the response to truckgusts. Fig. 13 shows measured acceleration at the center of the

Table 14Approximate stresses expected in critical members if the maximum sign area and span le

Sign truss Chord

Scaled model stress [MPa (ksi)] Scaled extrapolated stress [MPa (ks

II-C-A 77.1 (11.2) 112.2 (16.3)I-A 74.4 (10.8) 119.1 (17.3)II-A 86.1 (12.5) 130.8 (19.0)III-A 66.7 (9.67) 82.4 (11.9)

Fig. 13. Type III-A horizontal and vertical accelerations from truck gusts (le

structure for a 500-s example period of testing. Table 15 givesthe maximum acceleration and maximum stress recorded in achord member and a horizontal diagonal; note that these are peakstresses and not stress ranges. More detailed statistical analysis ofresults from all the truck gust measurements could be used to helpevaluate the cumulative effects of this sort of fatigue loading onsign trusses, but is outside the scope of this paper and will be pre-sented later elsewhere. For the Type III-A sign truss, the averagepeak acceleration due to truck gusts was approximately 300 mg,corresponding to a chord stress of 1.6 MPa (0.23 ksi) and an HDstress of 1.2 MPa (0.17 ksi).

5.5. Fatigue

Highway sign structural support fatigue design specifications [1]are based on a yearly mean wind speed of 5 m/s (11.2 mph). To com-pare results obtained in the field to the AASHTO design formulas, thevalue of maximum stress in each member was determined frommeasured strains occurring when this yearly wind speed actsnormal to the truss (as noted earlier, mean stress values corre-sponding to such maximums at this wind speed may only be aboutone-third as much). These stress values can then be compared to theConstant Amplitude Fatigue Limits (CAFL) for each member. Toshow the possible absolute maximum combined effect of a meanwind speed of 5 m/s (11.2 mph) and truck gusts, maximum stressesfrom each event have also been superimposed. Table 16 shows theseuncombined and combined stresses measured in each field-testedmember of the Type III-A sign structure, and for reference comparesthem to the CAFL values.

In this truss, the horizontal diagonal member (located on thebottom panel of the truss) did experience some stress ranges largerthan the CAFL, under both a 5 m/s (11.2 mph) wind speed andtruck gusts acting separately. The absolute maximum truck guststress used in this calculation for the HD is 5.8 MPa (0.84 ksi),

ngth are utilized.

Horizontal diagonal

i)] Scaled model stress [MPa (ksi)] Scaled extrapolated stress [MPa (ksi)]

15.0 (2.18) 17.6 (2.55)– –41.4 (6.02) 42.6 (6.21)62.3 (9.04) 49.6 (7.2)

ft) and detail of horizontal and vertical accelerations – truck 6 (right).

Table 15Type III-A acceleration and member stresses due to truck gusts.

Truck number Truck type Peak values

Horizontal acceleration (g) Chord stress [MPa (ksi)] Horizontal diagonal stress (MPa (ksi))

1 Typical with deflector 0.027 1.15 (0.167) 1.07 (0.155)2 Typical with deflector 0.030 1.49 (0.216) 1.16 (0.168)3 Typical w/o deflector 0.036 1.90 (0.275) 1.17 (0.170)4 Typical w/o deflector 0.028 1.77 (0.257) 1.39 (0.201)5 Typical w/o deflector 0.020 1.12 (0.162) 0.903 (0.131)6 Typical with deflector 0.045 2.63 (0.381) 1.93 (0.280)

Table 16Measured stress ranges compared to CAFL values.

Member Max. response to 5 m/s (11.2 mph) wind[MPa (ksi)]

Max. response to truck gust[MPa (ksi)]

5 m/s (11.2 mph) Wind + truckgust[MPa (ksi)]

CAFL [MPa(ksi)]

Chord 1 2.6 (0.37) 4.5 (0.66) 7.1 (1.03) 13 (1.9)Chord 2 2.6 (0.38) 6.1 (0.88) 8.7 (1.26) 13 (1.9)HD 3.2 (0.47) 5.8 (0.84) 9.0 (1.31) 3.0 (0.44)V1 2.0 (0.29) 0.55 (0.08) 2.6 (0.37) 3.0 (0.44)V2 1.0 (0.15) 0.28 (0.04) 1.3 (0.19) 3.0 (0.44)

J.A. Rice et al. / Engineering Structures 34 (2012) 173–186 185

which is considerably higher than the average peak stress due totruck gusts in the HD member of 1.2 MPa (0.17 ksi). However,these HD stress values could be even higher if the maximumspan/sign area were utilized, as previously discussed. The fact thatthe maximum truck gust stresses can exceed those due to the de-sign yearly mean wind speed is also noteworthy. On the otherhand, the version of the AASHTO specifications with which theseOSS were designed does not even require that fatigue be checkedfor span type structures, but rather only for cantilever structures,and as noted earlier the whole approach of using a reference pres-sure to check stresses versus the CAFL may not be the best methodin terms of fatigue analysis for design.

6. Discussion

Evaluation of the five sign truss structures confirmed that theyare in general compliance with the AASHTO specifications andIDOT standards that they have been designed to meet. However,the AASHTO design provisions themselves may not be completelyadequate in some areas, as discussed below. Sign truss structuresconstructed based on current IDOT designs do not appear to pres-ent significant risks of premature damage or failure. In fact, IDOTdesign calculations even make certain simplifications to their de-sign procedures that are somewhat more conservative than theAASHTO specifications.

When field-measured stresses were projected up to the staticdesign wind speed of 40.2 m/s (90 mph), two of the sign structures(the Types I-A and II-A) had members that did not satisfy a pure ax-ial allowable tension stress limit. The cantilever (II-C-A) structureand the Type III-A structure had projected stresses below theallowable stress limits (but both of these structures had consider-ably smaller sign areas than allowed by IDOT design tables). Stres-ses extrapolated from the measured values in the Type I-A andType II-A structures, with higher percentages of the maximumallowable sign areas, approach 138 MPa (20 ksi), which is the min-imum yield stress in the weld heat affected zone [1]. Those stresseshave a significant bending (stress gradient) component, so theoverstress is in a small area of the cross-section. Therefore,although the projected stress may exceed the allowable stress in

a small region of the chord member, safety does not appear to bean issue with the existing sign areas; utilizing the maximum allow-able sign areas, however, could result in larger stress levels.

The possibility of these overstresses is likely due to a combina-tion of factors. First, aspect ratios for IDOT overhead highway signsare in the range of 1.0–1.5, for which the drag coefficients wouldvary between 1.12 and 1.16 according to the AASHTO specifica-tions (2001). According to ASCE 7-05, the drag coefficients thatshould be used for IDOT signs that are part of this project are inthe range of 1.75–1.80 [3], more than 50% greater than those givenby the 2001 AASHTO specifications. IDOT calculations show that aconstant drag coefficient value of 1.20 was used in all their signtruss designs, which serves to mitigate this effect to a modestextent.

Another factor that affected the projected stresses to a smallerdegree is that current design procedures do not explicitly accountfor the vibration of a sign structure at ultimate wind loading(although this may have been a factor implicitly considered whenthe gust factor was developed). At a 40.2 m/s (90 mph) wind speed,field data suggest that the stress due to dynamic response will onlybe about 5% larger than the equivalent static design load; this is amean value, but the coefficient of variation is expected to be smallfor this wind speed.

And finally, IDOT design procedures conservatively assume thata 62 kPa (10 psf) uniform stress acts on the projected vertical crosssection area over regions where the sign is not present and a pres-sure of 1.4 kPa (30 psf) acts on the sign panels [when the actualcalculated pressure from Eq. (1) is 1.19 kPa (24.9 psf)]. These fac-tors (plus using Cd = 1.20) are the primary reasons why the pro-jected stresses are actually only a little bit larger than theallowable values.

AASHTO specifications have for some time required that canti-lever sign structures be designed against fatigue loads; however,simply supported OSS trusses have not typically been required tobe designed for fatigue. This study briefly evaluated fatigue effectsin all the sign truss structures investigated, for consideration byIDOT. Stress ranges experienced by selected truss members weremeasured under wind loads and truck gusts. Therefore, the fatiguestress range in each member could be evaluated considering a 5 m/s (11.2 mph) yearly mean wind speed and/or simultaneous truck

186 J.A. Rice et al. / Engineering Structures 34 (2012) 173–186

gust(s). Even maximum fatigue stress ranges measured in thechords of all the structures were lower than the allowable designstress range of 13.1 MPa (1.9 ksi), while the maximum fatiguestress ranges measured in web members were almost all smallerthan the allowable stress range of 3.0 MPa (0.44 ksi). If signsapproaching the maximum allowable area were placed on thesestructures, it is projected that the maximum stress ranges experi-enced by the horizontal diagonal members (but perhaps not themean stress ranges) would occasionally exceed the constant ampli-tude fatigue limit (CAFL). This is most likely to occur under thesimultaneous action of moderate wind in conjunction with largertruck gusts, a loading condition that is not currently mandatedfor design consideration by even the AASHTO specifications (andthat is quite conservative since peak vibration response to thewind, rather than average response, and peak vibration responseto truck gusts, would have to occur nearly simultaneously to pro-duce such a significant stress range in a member).

Given all the factors affecting the design (and actual) stresses insign truss members (and their corresponding measured stressranges), it is probably not practical to provide an accurate and con-sistent design equation for fatigue stresses within the current sim-plified reference pressure and CAFL framework. The sign structuredesigns developed and currently used by IDOT may be slightlyoverstressed with respect to fatigue in the horizontal diagonalmembers, especially if the maximum allowable sign areas are uti-lized. This is of particular note since AASHTO has not typically re-quired that OSS trusses be designed for fatigue. Additional fatigueassessments of these structures using more rational approachesthat can account for the cumulative effects of wind at variousvelocities and truck gust variability may be warranted.

7. Summary and conclusions

Using a combined analytical and full-scale experimentalapproach, four representative IDOT sign trusses have been evalu-ated to assess their AASHTO design load capacity and fatigue resis-tance. The experimental aspect of the study utilized a variety ofinstrumentation to capture the structural response (includingdamping) of the trusses under manual excitation, wind loading,and truck gust excitation. Analytical models were created for eachof the trusses and adjusted based on experimental results to mostaccurately represent the truss behavior. These models were thenutilized to assess the response of the trusses at design wind loads.Results indicate that there may be some cases where the trusseshave been designed such that modest overstresses could be possi-ble with respect to both static design wind loads and fatigue, anissue that is especially critical as the sign areas approach the max-imum allowed by IDOT specifications. The primary conclusion isthat if certain design load effects are updated based on the mostrecent suggested drag coefficients for this type of elevated signpanel, much of the overstress can be avoided. Additionally, resultsof this study suggest that OSS trusses may need to be evaluated for

fatigue, regardless of which AASHTO specifications (or moreadvanced approaches) are used.

Acknowledgments

This work was supported by the Illinois Department of Trans-portation (IDOT) as part of IDOT Contract No. IHR-R37 (Evaluationof Aluminum Highway Sign Truss Designs and Standards for Windand Truck Gust Loadings). The contents of this paper reflect theview of the writers, who are responsible for the facts and accuracyof the data presented herein. The contents do not necessarily re-flect the official views or policies of IDOT or the Federal HighwayAdministration, and this paper does not constitute a standard,specification, or regulation.

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