Bridges for Service Life Beyond 100 Years: Service Limit State Design (2014)

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217 In Chapter 6, various articles of AASHTO LRFD were identi- fied that would need to be modified to implement the calibrated SLS resulting from this research. This chapter contains the suggested modifications formatted in a form suitable for consideration by the affected technical committees that could be potential AASHTO Highway Subcommittee on Bridges and Structures agenda items. Excerpted material is used by permission of the American Association of State Highway and Transportation Officials. Since the various SLS revisions are independent of each other and could be implemented indi- vidually, the suggested provisions are presented in separate subsections for each SLS. The article numbering system used in AASHTO LRFD has been preserved. The proposed revisions are underlined and deletions are shown as strikethrough. C h a p t e r 7 Proposed Changes to AASHTO LRFD

218 7.1 Foundation Deformations – Service I 7.1.1 Proposed Revisions to Section 3 3.4—LOAD FACTORS AND COMBINATIONS 3.4.1—Load Factors and Load Combinations • • • • C3.4.1 • Service I—Load combination relating to the normal operational use of the bridge with a 55 mph wind and all loads taken at their nominal values. Also related to deflection control in buried metal structures, tunnel liner plate, and thermoplastic pipe, to control crack width in reinforced concrete structures, and for transverse analysis relating to tension in concrete segmental girders. This load combination should also be used for the investigation of slope stability, and settlement of foundations. Compression in prestressed concrete components and tension in prestressed bent caps are investigated using this load combination. Service III is used to investigate tensile stresses in prestressed concrete components. • Service II—Load combination intended to control yielding of steel structures and slip of slip-critical connections due to vehicular live load. This load combination corresponds to the overload provision for steel structures in past editions of the AASHTO Specifications, and it is applicable only to steel structures. From the point of view of load level, this combination is approximately halfway between that used for Service I and Strength I Limit States. • Service III—Load combination for longitudinal analysis relating to tension in prestressed concrete superstructures with the objective of crack control and to principal tension in the webs of segmental concrete girders. The live load specified in these specifications reflects, among other things, current exclusion weight limits mandated by various jurisdictions. Vehicles permitted under these limits have been in service for many years prior to 1993. For longitudinal loading, there is no nationwide physical evidence that these vehicles have caused cracking in existing prestressed concrete components. The statistical significance of the 0.80 factor on live load is that the event is expected to occur about once a year for bridges with two traffic lanes, less often for bridges with more than two traffic lanes, and about once a day for bridges with a single traffic lane. Service I should be used for checking tension related to transverse analysis of concrete segmental girders. The principal tensile stress check is introduced in order to verify the adequacy of webs of segmental concrete girder bridges for longitudinal shear and torsion. • Service IV—Load combination relating only to tension in prestressed concrete columns with the objective of crack control. The 0.70 factor on wind represents an 84 mph wind. This should result in zero tension in prestressed concrete columns for ten-year mean reoccurrence winds. The prestressed concrete columns must still meet strength requirements as set forth in Load Combination Strength III in Article 3.4.1. It is not recommended that thermal gradient be combined with high wind forces. Superstructure

219 expansion forces are included. • • • • The evaluation of overall stability of retained fills, as well as earth slopes with or without a shallow or deep foundation unit should be investigated at the service limit state based on the Service I Load Combination and an appropriate resistance factor as specified in Article 11.5.6 and Article 11.6.2.3. The investigation of foundation settlement shall proceed using the provisions of Article 10.6.2.4 using the load factor, γSE, specified in Table 3.4.1-4. For structural plate box structures complying with the provisions of Article 12.9, the live load factor for the vehicular live loads LL and IM shall be taken as 2.0. Applying these criteria for the evaluation of the sliding resistance of walls: • The vertical earth load on the rear of a cantilevered retaining wall would be multiplied by γpmin (1.00) and the weight of the structure would be multiplied by γpmin (0.90) because these forces result in an increase in the contact stress (and shear strength) at the base of the wall and foundation. • The horizontal earth load on a cantilevered retaining wall would be multiplied by γpmax (1.50) for an active earth pressure distribution because the force results in a more critical sliding force at the base of the wall. Similarly, the values of γpmax for structure weight (1.25), vertical earth load (1.35) and horizontal active earth pressure (1.50) would represent the critical load combination for an evaluation of foundation bearing resistance. Water load and friction are included in all strength load combinations at their respective nominal values. For creep and shrinkage, the specified nominal values should be used. For friction, settlement, and water loads, both minimum and maximum values need to be investigated to produce extreme load combinations. The load factor for temperature gradient, γTG, should be considered on a project-specific basis. In lieu of project-specific information to the contrary, γTG may be taken as: • 0.0 at the strength and extreme event limit states, • 1.0 at the service limit state when live load is not considered, and • 0.50 at the service limit state when live load is considered. The load factor for temperature gradient should be determined on the basis of the: • Type of structure, and • Limit state being investigated. Open girder construction and multiple steel box girders have traditionally, but perhaps not necessarily correctly, been designed without consideration of temperature gradient, i.e., γTG = 0.0.

220 The effects of the foundation deformation on the bridge superstructure, retaining walls, or other load bearing structures shall be evaluated at applicable strength and service limit states using the provisions of Article 10.5.2.2 and the settlement load factor (γSE) specified in Table 3.4.1-4. The load factor for settlement, γSE, should be considered on a project-specific basis. In lieu of project- specific information to the contrary, γSE, may be taken as 1.0. Load combinations which include settlement shall also be applied without settlement. For segmentally constructed bridges, the following combination shall be investigated at the service limit state: DC DW EH EV ES WA CR SH TG EL PS+ + + + + + + + + + (3.4.1-2) Methods for estimation of settlement based on local geologic conditions and calibration may be used subject to approval from the Owner. Calibration of local methods should be based on processes as described in SHRP 2 R19B program report (Kulicki et al., 2013). The value of γSE=1.25 for soil-structure interaction methods in Table 3.4.1-4 for estimation of lateral deformations has been established based on judgment at this time. Table 3.4.1-1—Load Combinations and Load Factors Load Combination Limit State DC DD DW EH EV ES EL PS CR SH LL IM CE BR PL LS WA WS WL FR TU TG SE Use One of These at a Time EQ BL IC CT CV Strength I (unless noted) γp 1.75 1.00 — — 1.00 0.50/1.20 γTG γSE — — — — — Strength II γp 1.35 1.00 — — 1.00 0.50/1.20 γTG γSE — — — — — Strength III γp — 1.00 1.4 0 — 1.00 0.50/1.20 γTG γSE — — — — — Strength IV γp — 1.00 — — 1.00 0.50/1.20 — — — — — — — Strength V γp 1.35 1.00 0.4 0 1.0 1.00 0.50/1.20 γTG γSE — — — — — Extreme Event I γp γEQ 1.00 — — 1.00 — — — 1.00 — — — — Extreme Event II γp 0.50 1.00 — — 1.00 — — — — 1.00 1.00 1.00 1.00 Service I 1.00 1.00 1.00 0.3 0 1.0 1.00 1.00/1.20 γTG γSE — — — — — Service II 1.00 1.30 1.00 — — 1.00 1.00/1.20 — — — — — — — Service III 1.00 0.80 1.00 — — 1.00 1.00/1.20 γTG γSE — — — — — Service IV 1.00 — 1.00 0.7 0 — 1.00 1.00/1.20 — 1.0 — — — — — Fatigue I— LL, IM & CE only — 1.50 — — — — — — — — — — — — Fatigue II— LL, IM & CE only — 0.75 — — — — — — — — — — — — • • • •

221 Table 3.4.1-4—Load Factors for Permanent Loads Due to Foundation Deformations, γSE Foundation Deformation and Deformation Estimation Method SE Immediate Settlement • Hough method 1.00 • Schmertmann method 1.25 • Local method * Consolidation settlement 1.00 Lateral Deformation • Soil-structure interaction method (P-y or Strain Wedge) 1.25 • Local method * *To be determined by the owner based on local geologic conditions and calibration using a target reliability index of 0.50 for Service I limit state. • • • • 3.4.2—Load Factors for Construction Loads 3.4.2.2—Evaluation of Deflection at the Service Limit State In the absence of special provisions to the contrary, where evaluation of construction deflections are required by the contract documents, Load Combination Service I shall apply. Construction dead loads shall be considered as part of the permanent load and construction transient loads considered part of the live load. The associated permitted deflections shall be included in the contract documents. Refer to Article 3.4.1 for evaluation of foundation deformations due to construction loads.

222 7.1.2 Proposed Revisions to Section 10 10.3—NOTATION Ad = angular distortion (10.5.2) Adm = modified angular distortion (10.5.2) C1 = correction factor to incorporate the effect of strain relief due to embedment (10.6.2.4.2b) C2 = correction factor to incorporate time-dependent (creep) increase in settlement for t (years) after construction (10.6.2.4.2b) E = modulus of elasticity of pile material (ksi) (10.7.3.8.2); elastic modulus of layer i based on guidance provided in Table C10.4.6.3-1 Iz = strain influence factor from Figure 10.6.2.4.2c-1a Ls = bridge span length over which Ad and Adm are computed (10.5.2) Sd = differential settlement between two bridge support elements spaced at a distance of Ls (ft) (10.5.2.2) Sta = total foundation settlement using all applicable loads in the Service I load combination (ft) (10.5.2) Stp = total foundation settlement using all applicable loads prior to construction of bridge superstructure in the Service I load combination (ft) (10.5.2.2) Str = relevant total settlement defined as Sta – Stp (10.5.2.2) X = width or smallest dimension of pile group (ft) (10.7.3.9); a factor used to determine the value of elastic modulus (10.6.2.4.2b) = load factor for settlement (10.5.2.2) ∆p = net uniform applied stress (load intensity) at the foundation depth (Figure 10.6.2.4.2c-1b) 10.5—LIMIT STATES AND RESISTANCE FACTORS 10.5.1—General The limit states shall be as specified in Article 1.3.2; foundation-specific provisions are contained in this Section. Foundations shall be proportioned so that the factored resistance is not less than the effects of the factored loads specified in Section 3. 10.5.2—Service Limit States 10.5.2.1—General Foundation design at the service limit state shall include: • Settlements, C10.5.2.1 In bridges where the superstructure and substructure are not integrated, settlement corrections can be made by jacking and shimming bearings. Article 2.5.2.3 requires jacking provisions for these bridges. • Horizontal movements, • Overall stability, and • Scour at the design flood. Consideration of foundation movements shall be based upon structure tolerance to total and differential movements, rideability and economy. Foundation movements shall include all movement from settlement, horizontal movement, and rotation. Bearing resistance estimated using the presumptive allowable bearing pressure for spread footings, if used, shall be applied only to address the service limit state. The cost of limiting foundation movements should be compared with the cost of designing the superstructure so that it can tolerate larger movements or of correcting the consequences of movements through maintenance to determine minimum lifetime cost. The Owner may establish more stringent criteria. The foundation movements should be translated to the deck elevation to evaluate the effect of such movements on the superstructure. In this process, deformations of the substructure, i.e., elements between foundation and superstructure, should be added to foundation deformations as appropriate.

223 The foundations for retaining walls and other load bearing structures such as tunnels may also be evaluated using the provisions of this Article. The design flood for scour is defined in Article 2.6.4.4.2, and is specified in Article 3.7.5 as applicable at the service limit state. Presumptive bearing pressures were developed for use with working stress design. These values may be used for preliminary sizing of foundations, but should generally not be used for final design. If used for final design, presumptive values are only applicable at service limit states. 10.5.2.2—Tolerable Movements and Movement Criteria 10.5.2.2.1—General Foundation movement criteria shall be consistent with the function and type of structure, anticipated service life, and consequences of unacceptable movements on structure performance. Foundation movement shall include vertical, horizontal, and rotational movements. The tolerable movement criteria shall be established by either empirical procedures or structural analyses, or by consideration of both. Foundation settlement shall be investigated using all applicable loads in the Service I Load Combination specified in Table 3.4.1-1. Transient loads may be omitted from settlement analyses for foundations bearing on or in cohesive soil deposits that are subject to time-dependent consolidation settlements. All applicable service limit state load combinations in Table 3.4.1-1 shall be used for evaluating horizontal movement and rotation of foundations. All foundation deformation evaluations shall be based on the geomaterial information obtained in accordance with Article 10.4. The following steps shall be followed to estimate a practical value of angular distortion of the superstructure based on foundation settlement; a similar approach can be applied and is recommended for evaluation of horizontal movement and rotation of foundations: 1. Compute total foundation settlement at each support element using an Owner approved method for the assumed foundation type (e.g., spread footings, driven piles, drilled shafts, etc.) as follows: a. Determine the total foundation settlement, Sta, using all applicable loads in the Service I load combination. b. Determine the total foundation settlement, Stp, prior to construction of bridge superstructure. This settlement would generally be as a result of all applicable substructure loads computed in accordance with Service I load combination. c. Determine relevant total settlement, Str as Str = Sta – Stp. C10.5.2.2.1 Experience has shown that bridges can and often do accommodate more movement and/or rotation than traditionally allowed or anticipated in design. Creep, relaxation, and redistribution of force effects accommodate these movements. Some studies have been made to synthesize apparent response. These studies indicate that angular distortions between adjacent foundations greater than 0.008 radians in simple spans and 0.004 radians in continuous spans should not be permitted in settlement criteria (Moulton et al., 1985; DiMillio, 1982; Barker et al., 1991; Samtani et al. 2010). Other angular distortion limits may be appropriate after consideration of: • cost of mitigation through larger foundations, realignment or surcharge, • rideability, • vertical clearance • tolerable limits of deformation of other structures associated with a bridge, e.g., approach slabs, wingwalls, pavement structures, drainage grades, utilities on the bridge, etc. • roadway drainage • aesthetics, and • safety. The bridge engineer shall add deformations from the substructure (elements between foundation and superstructure) as appropriate in evaluation of angular distortions at the deck elevation. While the angular distortion is generally applied in the longitudinal direction of a bridge, similar analyses should be performed in transverse direction based on consideration of bridge width and stiffness. For all bridges, stiffness should be appropriate to the considered limit state. Similarly, the effects of continuity with the substructure should be considered. In assessing the structural implications of foundation deformations of concrete bridges, the determination of the stiffness of

224 the bridge components should consider the effects of cracking, creep, and other inelastic responses Example: In Figure C10.5.2.2-1, a hypothetical 4- span bridge structure with span lengths, Ls1, Ls2, Ls3 and Ls4. The relevant total settlement, Str, is computed at each support element and the profile of Str along the bridge is shown by the solid line. In this example, Str-A1 < Str-P1 > Str-P2 < Str-P3 < Str-A2. The Str profile assumed for computation of the angular distortion, Ad, for each span is represented by the dashed lines. 2. At a given support element assume that the actual relevant settlement could be as large as the value calculated by the chosen method. At the same time, assume that the settlement of an adjacent support element could be zero instead of the relevant settlement value calculated by the same chosen method. Thus, differential settlement, Sd, within a given bridge span is equal to the larger of the relevant settlement at each of two supports of a bridge span. Compute angular distortion, Ad, as the ratio of the differential settlement, Sd, to the span length, Ls. Express Ad value in radians. Figure C10.5.2.2-1—Example for Computing Angular Distortion, Ad, Based on Relevant Total Settlement, Str, along a hypothetical 4-span Bridge (Modified after Samtani, et al., 2010) 3. Compute modified angular distortion, Adm, by multiplying the angular distortion value from Step 2 with the γSE values for settlement in Table 3.4.1-4 based on the method used for computing the total settlement value. 4. Compare the Adm value with owner specified angular distortion criteria. If owner specified criteria is not available then use 0.008 radians for the case of simple spans and 0.004 radians for the case of continuous spans as the limiting angular distortions. 5. Evaluate the structural ramifications of the computed angular distortions that are within acceptable limits as per Step 4. Modify foundation design as appropriate based on structural ramifications. The above procedure shall also be used for the cases where foundations of various support elements are proportioned for equal total settlement because the prediction of settlements from any given method is uncertain by itself. The angular distortion, Ad, within each span is as follows: Ad1 = Str-P1/Ls1; Ad1 = Str-P1/Ls2; Ad3 = Str-P3/Ls3; and Ad4 = Str-A2/Ls4. Express Ad value in radians. Multiply the Ad values with appropriate γSE as per Step 3. 10.5.2.2.2—Lateral Deformations Using a procedure similar to settlement evaluation specified in Article 10.5.2.2.1, lateral (horizontal) movement at foundation level shall also be evaluated. Horizontal movement criteria should be established at the top of the foundation based on the tolerance of the structure to lateral movement, with consideration of the column length and stiffness. Table 3.4.1-4 provides C10.5.2.2.2 Rotation movements should be evaluated at the top of the substructure unit in plan location and at the deck elevation. Tolerance of the superstructure to lateral movement will depend on bridge seat or joint widths, bearing type(s), structure type, and load distribution effects.

225 values of γSE for lateral deformations. 10.5.2.2.3—Walls The procedure for computing angular distortions shall also be applied for evaluating angular distortions along and transverse to retaining walls as well as the junction of the approach walls to abutment walls. The angular distortion values along a retaining wall can be used to select an appropriate wall type, e.g., MSE walls can tolerate larger angular distortions compared to cast- in-place walls 10.5.2.3—Overall Stability The evaluation of overall stability of earth slopes with or without a foundation unit shall be investigated at the service limit state as specified in Article 11.6.2.3. 10.5.2.4—Abutment Transitions Vertical and horizontal movements caused by embankment loads behind bridge abutments shall be investigated. C10.5.2.4 Settlement of foundation soils induced by embankment loads can result in excessive movements of substructure elements. Both short and long term settlement potential should be considered. Settlement of improperly placed or compacted backfill behind abutments can cause poor rideability and a possibly dangerous bump at the end of the bridge. Guidance for proper detailing and material requirements for abutment backfill is provided in Cheney and Chassie Samtani and Nowatzki (20006). Lateral earth pressure behind and/or lateral squeeze below abutments can also contribute to lateral movement of abutments and should be investigated, if applicable. 10.6.2—Service Limit State Design 10.6.2.1—General • • • • C10.6.2.1 10.6.2.4—Settlement Analyses 10.6.2.4.1—General Foundation settlements should be estimated using computational methods based on the results of laboratory or insitu testing, or both. The soil parameters used in the computations should be chosen to reflect the loading history of the ground, the construction sequence, and the effects of soil layering. Both total and differential settlements, including time dependant effects, shall be considered. Total settlement, including elastic, consolidation, C10.6.2.4.1 Elastic, or immediate, settlement is the instantaneous deformation of the soil mass that occurs as the soil is loaded. The magnitude of elastic settlement is estimated as a function of the applied stress beneath a footing or embankment. Elastic settlement is usually small and neglected in design, but where settlement is critical, it is the most important deformation consideration in cohesionless soil deposits and for footings bearing on rock. For footings located on over-

226 and secondary components may be taken as: t e c sS S S S= + + (10.6.2.4.1-1) where: Se = elastic settlement (ft) Sc = primary consolidation settlement (ft) Ss = secondary settlement (ft) consolidated clays, the magnitude of elastic settlement is not necessarily small and should be checked. In a nearly saturated or saturated cohesive soil, the pore water pressure initially carries the applied stress. As pore water is forced from the voids in the soil by the applied load, the load is transferred to the soil skeleton. Consolidation settlement is the gradual compression of the soil skeleton as the pore water is forced from the voids in the soil. Consolidation settlement is the most important deformation consideration in cohesive soil deposits that possess sufficient strength to safely support a spread footing. While consolidation settlement can occur in saturated cohesionless soils, the consolidation occurs quickly and is normally not distinguishable from the elastic settlement. Secondary settlement, or creep, occurs as a result of the plastic deformation of the soil skeleton under a constant effective stress. Secondary settlement is of principal concern in highly plastic or organic soil deposits. Such deposits are normally so obviously weak and soft as to preclude consideration of bearing a spread footing on such materials. The principal deformation component for footings on rock is elastic settlement, unless the rock or included discontinuities exhibit noticeable time-dependent behavior. To avoid overestimation, relevant settlements should be evaluated using the construction point concept noted in Samtani et al. (2010). The effect of settlement on superstructure shall be evaluated based on Article 10.5.2.2. The effects of the zone of stress influence, or vertical stress distribution, beneath a footing shall be considered in estimating the settlement of the footing. Spread footings bearing on a layered profile consisting of a combination of cohesive soil, cohesionless soil and/or rock shall be evaluated using an appropriate settlement estimation procedure for each layer within the zone of influence of induced stress beneath the footing. The distribution of vertical stress increase below circular or square and long rectangular footings, i.e., where L > 5B, may be estimated using Figure 10.6.2.4.1-1. For guidance on vertical stress distribution for complex footing geometries, see Poulos and Davis (1974) or Lambe and Whitman (1969). Some methods used for estimating settlement of footings on sand include an integral method to account for the effects of vertical stress increase variations. For guidance regarding application of these procedures, see Gifford et al. (1987).

227 Figure 10.6.2.4.1-1—Boussinesq Vertical Stress Contours for Continuous and Square Footings Modified after Sowers (1979) 10.6.2.4.2—Settlement of Footings on Cohesionless Soils 10.6.2.4.2a—General C10.6.2.4.2a The settlement of spread footings bearing on cohesionless soil deposits shall be estimated as a function of effective footing width and shall consider the effects of footing geometry and soil and rock layering with depth. Although methods are recommended for the determination of settlement of cohesionless soils, experience has indicated that settlements can vary considerably in a construction site, and this variation may not be predicted by conventional calculations. Settlements of cohesionless soils occur rapidly, essentially as soon as the foundation is loaded. Therefore, the total settlement under the service loads may not be as important as the incremental settlement between intermediate load stages. For example, the total and differential settlement due to loads applied by columns and cross beams is generally less important than the total and differential settlements due to girder placement and casting of continuous concrete decks. Settlements of footings on cohesionless soils shall be estimated using elastic theory or empirical procedures. Generally conservative settlement estimates may be obtained using the elastic half-space procedure or the empirical method by Hough. Additional information regarding the accuracy of the methods described herein is provided in Gifford et al. (1987), and Kimmerling (2002) and Samtani and Notwazki (2006). This information, in combination with local experience and engineering judgment, should be used when determining the estimated settlement for a structure foundation, as there may be cases, such as attempting to build a structure grade high to account for the estimated settlement, when overestimating the settlement magnitude could be problematic. Details of other procedures can be found in

228 textbooks and engineering manuals, including: • Terzaghi and Peck (1967) • Sowers (1979) • U.S. Department of the Navy (1982) • D’Appolonia (Gifford et al., 1987)—This method includes consideration for over- consolidated sands. • Tomlinson (1986) • Gifford et al. (1987) 10.6.2.4.2b—Elastic Half-space Method The elastic half-space method assumes the footing is flexible and is supported on a homogeneous soil of infinite depth. The elastic settlement of spread footings, in feet, by the elastic half-space method shall be estimated as: ( )21 144 E β q A o S e s z  ′−  = ν (10.6.2.4.2b-1) where: qo = applied vertical stress (ksf) A′ = effective area of footing (ft2) Es = Young’s modulus of soil taken as specified in Article 10.4.6.3 if direct measurements of Es are not available from the results of in situ or laboratory tests (ksi) βz = shape factor taken as specified in Table 10.6.2.4.2b-1 (dim) ν = Poisson’s Ratio, taken as specified in Article 10.4.6.3 if direct measurements of ν are not available from the results of in situ or laboratory tests (dim) Unless Es varies significantly with depth, Es should be determined at a depth of about 1/2 to 2/3 of B below the footing, where B is the footing width. If the soil modulus varies significantly with depth, a weighted average value of Es should be used. C10.6.2.4.2b For general guidance regarding the estimation of elastic settlement of footings on sand, see Gifford et al. (1987), and Kimmerling (2002), and Samtani and Notwazki (2006). The stress distributions used to calculate elastic settlement assume the footing is flexible and supported on a homogeneous soil of infinite depth. The settlement below a flexible footing varies from a maximum near the center to a minimum at the edge equal to about 50 percent and 64 percent of the maximum for rectangular and circular footings, respectively. The settlement profile for rigid footings is assumed to be uniform across the width of the footing. Spread footings of the dimensions normally used for bridges are generally assumed to be rigid, although the actual performance will be somewhere between perfectly rigid and perfectly flexible, even for relatively thick concrete footings, due to stress redistribution and concrete creep. The accuracy of settlement estimates using elastic theory are strongly affected by the selection of soil modulus and the inherent assumptions of infinite elastic half space. Accurate estimates of soil moduli are difficult to obtain because the analyses are based on only a single value of soil modulus, and Young’s modulus varies with depth as a function of overburden stress. Therefore, in selecting an appropriate value for soil modulus, consideration should be given to the influence of soil layering, bedrock at a shallow depth, and adjacent footings. For footings with eccentric loads, the area, A′, should be computed based on reduced footing dimensions as specified in Article 10.6.1.3.

229 Table 10.6.2.4.2b-1—Elastic Shape and Rigidity Factors, EPRI (1983) L/B Flexible, βz (average) βz Rigid Circular 1.04 1.13 1 1.06 1.08 2 1.09 1.10 3 1.13 1.15 5 1.22 1.24 10 1.41 1.41 10.6.2.4.2c—Hough Method Estimation of spread footing settlement on cohesionless soils by the empirical Hough method shall be determined using Eqs. 10.6.2.4.2c-2 and 10.6.2.4.2c-3. SPT blow counts shall be corrected as specified in Article 10.4.6.2.4 for depth, i.e. overburden stress, before correlating the SPT blow counts to the bearing capacity index, C ′. 1 n e i i S H = = ∆∑ (10.6.2.4.2c-1) in which: 1 log o v o i c C H H ′σ + ∆σ ′ ′σ   ∆ =     (10.6.2.4.2c-2) where: n = number of soil layers within zone of stress influence of the footing ∆Hi = elastic settlement of layer i (ft) HC = initial height of layer i (ft) C′ = bearing capacity index from Figure 10.6.2.4.2c-1 (dim) C10.6.2.4.2c The Hough method was developed for normally consolidated cohesionless soils. The Hough method has several advantages over other methods used to estimate settlement in cohesionless soil deposits, including express consideration of soil layering and the zone of stress influence beneath a footing of finite size. The subsurface soil profile should be subdivided into layers based on stratigraphy to a depth of about three times the footing width. The maximum layer thickness should be about 10 ft. While Cheney and Chassie (2000), and Hough (1959), did not specifically state that the SPT N values should be corrected for hammer energy in addition to overburden pressure, due to the vintage of the original work, hammers that typically have an efficiency of approximately 60 percent were in general used to develop the empirical correlations contained in the method. If using SPT hammers with efficiencies that differ significantly from this 60 percent value, the N values should also be corrected for hammer energy, in effect requiring that N160 be used (Samtani and Nowatzki, 2006). Studies conducted by Gifford et al. (1987) and Samtani and Nowatzki (2006) indicate that Hough’s procedure is conservative and over-predicts settlement by a factor of 2 or more. Such conservatism may be acceptable for the evaluation of the settlement of embankments. However, in the case of shallow foundations such conservatism may lead to unnecessary use of costlier deep foundations in cases where shallow foundations may be viable. In Figure 10.6.2.4.2c-1, N1 shall be taken as N160, Standard Penetration Resistance, N (blows/ft), corrected for overburden pressure as specified in Article 10.4.6.2.4. σ′o = initial vertical effective stress at the midpoint of layer i (ksf) ∆σv = increase in vertical stress at the midpoint of layer i (ksf)

230 Figure 10.6.2.4.2c-1—Bearing Capacity Index versus Corrected SPT (Samtani and Nowatzki, 2006, after Hough, 1959) The Hough method is applicable to cohesionless soil deposits. The “Inorganic Silt” curve should generally not be applied to soils that exhibit plasticity because N-values in such soils are unreliable. The settlement characteristics of cohesive soils that exhibit plasticity should be investigated using undisturbed samples and laboratory consolidation tests as prescribed in Article 10.6.2.4.3. 10.6.2.4.2d—Schmertmann Method An estimate of the immediate settlement, Si, of spread footings may be made by using Eq. 10.6.2.4.2d-1 as proposed by Schmertmann, et al. (1978). n i 1 2 i i=1 S = C C Dp DH∑ (10.6.2.4.2d-1) in which: Δ zi c I H H XE  =     (10.6.2.4.2d-2) 5.0 p p 5.01C o1 ≥      ∆ −= (10.6.2.4.2d-3) ( )      += 1.0 yearst log2.01C 102 (10.6.2.4.2d-4) where: Iz = strain influence factor from Figure 10.6.2.4.2d- 1a. The dimension Bf represents the least lateral dimension of the footing after correction for eccentricities, i.e. use least lateral effective footing dimension. The strain influence factor is a function of depth and is obtained from the strain influence diagram. The strain influence diagram is easily constructed for the C10.6.2.4.2d To overcome the conservatism of the Hough method, use of a more rigorous procedure such as Schmertmann’s method (1978) may be used for shallow foundations. • Effect of lateral strain: Schmertmann method is based on the results of displacement measurements within sand masses loaded by model footings, as well as finite element analyses of deformations of materials with nonlinear stress-strain behavior that expressly incorporated Poisson’s ratio. Therefore, the effect of the lateral strain on the vertical strain is included in the strain influence factor diagrams. • Effect of preloading: The equations used in Schmertmann’s method are applicable to normally loaded sands. If the sand was pre-strained by previous loading, then the actual settlements will be overpredicted. Schmertmann, et al. (1970) and Holtz (1991) recommend a reduction in settlement after preloading or other means of compaction of half the predicted settlement. Alternatively, in case of preloaded soil deposits, the settlement can be computed by using the method proposed by D’Appolonia (1968, 1970), which includes explicit consideration of preloading. • C2 correction factor and applicability of the method: The time duration, t, in Eq. 10.6.2.4.2d-4 is set to 0.1 years to evaluate the settlement immediately after construction, i.e., C2 = 1. If long-

231 axisymmetric case (Lf/Bf = 1) and the plane strain case (Lf/Bf ≥ 10) as shown in Figure 10.6.2.4.2d-1a. The strain influence diagram for intermediate conditions can be determined by simple linear interpolation. n = number of soil layers within the zone of strain influence (strain influence diagram). ∆p = net uniform applied stress (load intensity) at the foundation depth (see Figure 10.6.2.4.2d-1b). E = elastic modulus of layer i based on guidance provided in Table C10.4.6.3-1. X = a factor used to determine the value of elastic modulus. If the value of elastic modulus is based on correlations with N160-values or qc from Table C10.4.6.3-1, then use X as follows. X = 1.25 for axisymmetric case (Lf/Bf = 1) X = 1.75 for plane strain case (Lf/Bf ≥ 10) Use interpolation for footings with 1 < Lf/Bf ≤ 10 If the value of elastic modulus is estimated based on the range of elastic moduli in Table C10.4.6.3-1 or other sources use X = 1.0. C1 = correction factor to incorporate the effect of strain relief due to embedment po = effective in-situ overburden stress at the foundation depth and ∆p is the net foundation pressure as shown in Figure 10.6.2.4.2d-1b C2 = correction factor to incorporate time-dependent (creep) increase in settlement for t (years) after construction where: (a) term creep deformation of the soil is suspected then an appropriate time duration, t, can be used in the computation of C2. Creep deformation is not the same as consolidation settlement. This factor can have an important influence on the reported settlement since it is included in Eq. 10.6.2.4.2d-1 as a multiplier. For example, the C2 factor for time durations of 0.1 yrs, 1 yr, 10 yrs and 50 yrs are 1.0, 1.2, 1.4 and 1.54, respectively. In cohesionless soils and unsaturated fine-grained cohesive soils with low plasticity, time durations of 0.1 yr and 1 yr, respectively, are generally appropriate and sufficient for cases of static loads. The C2 parameter shall not be used to estimate time- dependent consolidation settlements. Where consolidation settlement can occur within the depth of the strain distribution diagram, the magnitude of the consolidation settlement shall be estimated as per Article 10.6.2.4.3 and added to the immediate settlement of other layers within the strain distribution diagram where consolidation settlement may not occur.

232 (b) Figure 10.6.2.4.2d-1—(a) Simplified vertical strain influence factor distributions, (b) Explanation of pressure terms in equation for Izp (after Schmertmann, et al., 1978, Samtani and Notatzki, 2006). 10.6.2.4.2e—Local Method Methods based on local geologic conditions and calibration may be used subject to approval from the Owner. C10.6.2.4.2e Calibration of local methods should be based on processes as described in SHRP 2 R19B program report (Kulicki et al., 2013). 10.10—REFERENCES D'Appolonia, D. J., D'Appolonia, E. E., and Brissette, R. F. (1968). "Settlement of Spread Footings on Sand." American Society of Civil Engineers, Journal of the Soil Mechanics and Foundations Division, 94 (SM3), 735-760. D'Appolonia, D. J., D'Appolonia, E. E., and Brissette, R. F. (1970). Closure to discussions on "Settlement of Spread Footings on Sand." American Society of Civil Engineers, Journal of the Soil Mechanics and Foundations Division, 96 (SM2), 754-761. Holtz, R. D. 1991. “Stress Distribution and Settlement of Shallow Foundations.” In Foundation Engineering Handbook, 2nd Edition, H. Y. Fang, editor. Van Nostrand Reinhold Co., New York, NY, Chapter 5, pp. 166–185. Samtani, N. C., and Nowatzki, E. A. 2006. Soils and Foundations, FHWA NHI-06-088 and FHWA NHI 06-089, Federal Highway Administration, U.S. Department of Transportation, Washington, DC. Samtani, N. C., Nowatzki, E. A., and Mertz, D.R. 2010. Selection of Spread Footings on Soils to Support Highway Bridge Structures, FHWA-RC/TD-10-001, Federal Highway Administration, Resource Center, Matteson, IL Schmertmann, J. H. 1970. "Static Cone to Compute Static Settlement Over Sand." American Society of Civil Engineers, Journal of the Soil Mechanics and Foundations Division, 96(SM3), 1011-1043. Schmertmann, J. H., Hartman, J. P., and Brown, P. R. 1978. "Improved Strain Influence Factor Diagrams." American Society of Civil Engineers, Journal of the Geotechnical Engineering Division, 104 (No. GT8), 1131-1135.

233 7.2 Live Load Response 7.2.1 Proposed Revisions to Section 2 2.5.2.6—Deformations 2.5.2.6.1—General • • • • C2.5.2.6.1 • • • • 2.5.2.6.2—Criteria for Deflection Live Load Response The criteria in this Section shall be considered optional, except for the following: • The provisions for orthotropic decks shall be considered mandatory. • The provisions in Article 12.14.5.9 for precast reinforced concrete three-sided structures shall be considered mandatory. • Metal grid decks and other lightweight metal and concrete bridge decks shall be subject to the serviceability provisions of Article 9.5.2. In applying these criteria, the vehicular load shall include the dynamic load allowance. If an Owner chooses to invoke deflection control, the following principles may be applied: C2.5.2.6.2 These provisions permit, but do not encourage, the use of past practice for deflection control. Designers were permitted to exceed these limits at their discretion in the past. Calculated deflections of structures have often been found to be difficult to verify in the field due to numerous sources of stiffness not accounted for in calculations. Despite this, many Owners and designers have found comfort in the past requirements to limit the overall stiffness of bridges. The desire for continued availability of some guidance in this area, often stated during the development of these Specifications, has resulted in the retention of optional criteria, except for orthotropic decks, for which the criteria are required. Deflection criteria are also mandatory for lightweight decks comprised of metal and concrete, such as filled and partially filled grid decks, and unfilled grid decks composite with reinforced concrete slabs, as provided in Article 9.5.2. Additional guidance regarding deflection of steel bridges can be found in Wright and Walker (1971). Additional considerations and recommendations for deflection in timber bridge components are discussed in more detail in Chapters 7, 8, and 9 in Ritter (1990). • When investigating the maximum absolute deflection for straight girder systems, all design lanes should be loaded, and all supporting components should be assumed to deflect equally; • For curved steel box and I-girder systems, the deflection of each girder should be determined individually based on its response as part of a system; For a straight multibeam bridge, this is equivalent to saying that the distribution factor for deflection is equal to the number of lanes divided by the number of beams. For curved steel girder systems, the deflection limit is applied to each individual girder because the curvature causes each girder to deflect differently than the adjacent girder so that an average deflection has little meaning. For curved steel girder systems, the span used to compute the deflection limit should be taken as the arc girder length between bearings. • For composite design, the stiffness of the design cross-section used for the determination of deflection and frequency should include the entire width of the roadway and the structurally continuous portions of the railings, sidewalks, and median barriers; • For straight girder systems, the composite bending stiffness of an individual girder may be taken as the

234 stiffness determined as specified above, divided by the number of girders; • When investigating maximum relative displacements, the number and position of loaded lanes should be selected to provide the worst differential effect; • The live load portion of Load Combination Service I of Table 3.4.1-1 should be used, including the dynamic load allowance, IM; • The live load shall be taken from Article 3.6.1.3.2; • The provisions of Article 3.6.1.1.2 should apply; and • For skewed bridges, a right cross-section may be used, and for curved and curved skewed bridges, a radial cross-section may be used. In the absence of other criteria, the following deflection limits may be considered for steel, aluminum, and/or concrete vehicular bridges: should meet the criteria shown in Figure 2.5.2.6.1-1 for the anticipated level of pedestrian usage. Unless otherwise specified herein, the deflection and frequency may be calculated on a system or component basis using any recognized method of analysis. Figure 2.5.2.6.1-1—Criteria for Live Load Response (Used with permission of the Canadian Standards Association) • Vehicular load, general ............................. Span/800, • Vehicular and pedestrian loads ............... Span/1000, • Vehicular load on cantilever arms ............................ Span/300, and • Vehicular and pedestrian loads on cantilever arms... Span/375. For steel I-shaped beams and girders, and for steel box Frequency may be determined by refined analysis methods or by equations available in the literature if they apply to the girder or structure being analyzed.

235 and tub girders, the provisions of Articles 6.10.4.2 and 6.11.4, respectively, regarding the control of permanent deflections through flange stress controls, shall apply. For pedestrian bridges, i.e., bridges whose primary function is to carry pedestrians, bicyclists, equestrians, and light maintenance vehicles, the provisions of Section 5 of AASHTO’s LRFD Guide Specifications for the Design of Pedestrian Bridges shall apply. In the absence of other criteria, the following deflection limits may be considered for wood construction: • Vehicular and pedestrian loads ...........Span/425, and • Vehicular load on wood planks and panels (extreme relative deflection between adjacent edges) ...... 0.10 in. From a structural viewpoint, large deflections in wood components cause fasteners to loosen and brittle materials, such as asphalt pavement, to crack and break. In addition, members that sag below a level plane present a poor appearance and can give the public a perception of structural inadequacy. Deflections from moving vehicle loads also produce vertical movement and vibrations that annoy motorists and alarm pedestrians (Ritter, 1990). The following provisions shall apply to orthotropic plate decks: • Vehicular load on deck plate .................... Span/300, • Vehicular load on ribs of orthotropic metal decks Span/1000, and Excessive deformation can cause premature deterioration of the wearing surface and affect the performance of fasteners, but limits on the latter have not yet been established. The intent of the relative deflection criterion is to protect the wearing surface from debonding and fracturing due to excessive flexing of the deck. • Vehicular load on ribs of orthotropic metal decks (extreme relative deflection between adjacent ribs) 0.10 in. The 0.10-in. relative deflection limitation is tentative. 7.2.2 Proposed Revisions to Section 3 3.4—LOAD FACTORS AND COMBINATIONS 3.4.1—Load Factors and Load Combinations • • • • C3.4.1 • • • • • Service I—Load combination relating to the normal operational use of the bridge with a 55 mph wind and all loads taken at their nominal values. Also related to dynamic response of superstructures, deflection control in buried metal structures, tunnel liner plate, and thermoplastic pipe, to control crack width in reinforced concrete structures, and for transverse analysis relating to tension in concrete segmental girders. This load combination should also be used for the investigation of slope stability. Compression in prestressed concrete components and tension in prestressed bent caps are investigated using this load combination. Service III is used to investigate tensile stresses in prestressed concrete components. • Service II—Load combination intended to control yielding of steel structures and slip of slip-critical connections due to vehicular live load. This load combination corresponds to the overload provision for steel structures in past editions of the AASHTO Specifications, and it is applicable only to steel structures. From the point of view of load level, this combination is approximately halfway between that used for Service I and Strength I Limit States. • Service III—Load combination for longitudinal The live load specified in these specifications

236 analysis relating to tension in prestressed concrete superstructures with the objective of crack control and to principal tension in the webs of segmental concrete girders. reflects, among other things, current exclusion weight limits mandated by various jurisdictions. Vehicles permitted under these limits have been in service for many years prior to 1993. For longitudinal loading, there is no nationwide physical evidence that these vehicles have caused cracking in existing prestressed concrete components. The statistical significance of the 0.80 factor on live load is that the event is expected to occur about once a year for bridges with two traffic lanes, less often for bridges with more than two traffic lanes, and about once a day for bridges with a single traffic lane. Service I should be used for checking tension related to transverse analysis of concrete segmental girders. The principal tensile stress check is introduced in order to verify the adequacy of webs of segmental concrete girder bridges for longitudinal shear and torsion. • Service IV—Load combination relating only to tension in prestressed concrete columns with the objective of crack control. The 0.70 factor on wind represents an 84 mph wind. This should result in zero tension in prestressed concrete columns for ten-year mean reoccurrence winds. The prestressed concrete columns must still meet strength requirements as set forth in Load Combination Strength III in Article 3.4.1. It is not recommended that thermal gradient be combined with high wind forces. Superstructure expansion forces are included. • Service V—Load combination to be used to investigate deflection and vibration response under traffic in accordance with Article 2.5.2.6.2. Dead load is included in this load combination because mass is part of the required calculation of frequency. • • • •

237 Table 3.4.1-1—Load Combinations and Load Factors Load Combination Limit State DC DD DW EH EV ES EL PS CR SH LL IM CE BR PL LS WA WS W L FR TU TG SE Use One of These at a Time EQ BL IC CT CV Strength I (unless noted) γp 1.75 1.00 — — 1.00 0.50/1.20 γTG γSE — — — — — Strength II γp 1.35 1.00 — — 1.00 0.50/1.20 γTG γSE — — — — — Strength III γp — 1.00 1.40 — 1.00 0.50/1.20 γTG γSE — — — — — Strength IV γp — 1.00 — — 1.00 0.50/1.20 — — — — — — — Strength V γp 1.35 1.00 0.40 1.0 1.00 0.50/1.20 γTG γSE — — — — — Extreme Event I γp γEQ 1.00 — — 1.00 — — — 1.00 — — — — Extreme Event II γp 0.50 1.00 — — 1.00 — — — — 1.00 1.00 1.00 1.00 Service I 1.00 1.00 1.00 0.30 1.0 1.00 1.00/1.20 γTG γSE — — — — — Service II 1.00 1.30 1.00 — — 1.00 1.00/1.20 — — — — — — — Service III 1.00 0.80 1.00 — — 1.00 1.00/1.20 γTG γSE — — — — — Service IV 1.00 — 1.00 0.70 — 1.00 1.00/1.20 — 1.0 — — — — — Service V 1.00 1.50 — — — — — — — — — — — — Fatigue I— LL, IM & CE only — 1.50 — — — — — — — — — — — — Fatigue II— LL, IM & CE only — 0.75 — — — — — — — — — — — —

238 7.3 Premature Yielding and Slip of Bolts – Service II 7.3.1 Proposed Revisions to Section 3 3.4—LOAD FACTORS AND COMBINATIONS • • • • • Service I—Load combination relating to the normal operational use of the bridge with a 55 mph wind and all loads taken at their nominal values. Also related to deflection control in buried metal structures, tunnel liner plate, and thermoplastic pipe, to control crack width in reinforced concrete structures, and for transverse analysis relating to tension in concrete segmental girders. Relevant elements of this load combination should also be used for the investigation of slope stability. Compression in prestressed concrete components and tension in prestressed bent caps are investigated using this load combination. Service III is used to investigate tensile stresses in prestressed concrete components. • Service II—Load combination intended to control yielding of steel structures and slip of slip-critical connections due to vehicular live load. This load combination corresponds to the overload provision for steel structures in past editions of the AASHTO Specifications, and it is applicable only to steel structures. From the point of view of load level, this combination is approximately halfway between that used for Service I and Strength I Limit States. A recent evaluation of WIM data from 31 sites around the country indicated that the load level specified in Table 3.4.1-1 for this limit state could reasonably be expected to be exceeded less than once every six months on average. For structures with unique truck loading conditions, such as access roads to ports or industrial sites which might lead to a disproportionate number of permit loads, a site-specific increase in the load factor or number of loaded lanes should be considered. • Service III—Load combination for longitudinal analysis relating to tension in prestressed concrete superstructures with the objective of crack control and to principal tension in the webs of segmental concrete girders. The live load specified in these specifications reflects, among other things, current exclusion weight limits mandated by various jurisdictions. Vehicles permitted under these limits have been in service for many years prior to 1993. For longitudinal loading, there is no nationwide physical evidence that these vehicles have caused cracking in existing prestressed concrete components. The statistical significance of the 0.80 factor on live load is that the event is expected to occur about once a year for bridges with two traffic lanes, less often for bridges with more than two traffic lanes, and about once a day for bridges with a single traffic lane. Service I should be used for checking tension related to transverse analysis of concrete segmental girders. The principal tensile stress check is introduced in order to verify the adequacy of webs of segmental concrete girder bridges for longitudinal shear and torsion.

239 • Service IV—Load combination relating only to tension in prestressed concrete columns with the objective of crack control. The 0.70 factor on wind represents an 84 mph wind. This should result in zero tension in prestressed concrete columns for ten-year mean reoccurrence winds. The prestressed concrete columns must still meet strength requirements as set forth in Load Combination Strength III in Article 3.4.1. It is not recommended that thermal gradient be combined with high wind forces. Superstructure expansion forces are included. • • • •

240 7.4 Cracking of Prestressed Concrete – Currently Service III 7.4.1 Proposed Revisions to Section 3 3.4—LOAD FACTORS AND COMBINATIONS 3.4.1—Load Factors and Load Combinations The total factored force effect shall …………….. . ---. C3.4.1 The background for the load factors ……………….. . . Service I—Load combination relating to the normal operational use of the bridge with a 55 mph wind and all loads taken at their nominal values. Also related to deflection control in buried metal structures, tunnel liner plate, and thermoplastic pipe, to control crack width in reinforced concrete structures, and for transverse analysis relating to tension in concrete segmental girders. This load combination should also be used for the investigation of slope stability. Compression in prestressed concrete components and tension in prestressed bent caps are investigated using this load combination. Service III is used to investigate tensile stresses in prestressed concrete components. Service II—Load combination intended to control yielding of steel structures and slip of slip-critical connections due to vehicular live load. This load combination corresponds to the overload provision for steel structures in past editions of the AASHTO Specifications, and it is applicable only to steel structures. From the point of view of load level, this combination is approximately halfway between that used for Service I and Strength I Limit States. Service III—Load combination for longitudinal analysis relating to tension in prestressed concrete superstructures with the objective of crack control and to principal tension in the webs of segmental concrete girders. Prior to 2014, the longitudinal analysis relating to tension in prestressed concrete superstructures was investigated using a load factor for live load of 0.8. The live load specified in these specifications This load factor reflecteds, among other things, current exclusion weight limits mandated by various jurisdictions at the time of the development of the specifications in 1993. Vehicles permitted under these limits have been in service for many years prior to 1993. It was concluded at that time that, for longitudinal loading, there is no nationwide physical evidence that these vehicles have caused cracking in existing prestressed concrete components. The 0.8 load factor was applied regardless of the method used for determining the loss of prestressing. The statistical significance of the 0.80 factor on live load is that the event is expected to occur about once a year for bridges with two traffic lanes, less often for bridges with more than two traffic lanes, and about once a day for bridges with a single traffic lane. The calibration of the service limit states for concrete components (Wassef et. al. 2014) concluded that typical components designed using the Refined Estimates of Time-Dependent Losses method incorporated in the specifications in 2005 have a lower reliability index against flexural cracking in prestressed components than components designed using the prestress loss calculation method specified prior to 2005. For components designed using the currently-specified methods for instantaneous prestressing losses and the

241 currently-specified Refined Estimates of Time- Dependent Losses method, an increase in the load factor for live load from 0.8 to 1.0 was required to maintain the level of reliability against cracking of prestressed concrete components inherent in the system. Components which design satisfies all of the following conditions: • A refined time step method is used for calculating the time-dependent prestressing losses • The section properties are determined based on the concrete gross section, and, • The force in prestressing steel is determined without taking advantage of the elastic gain, were not affected by the changes in the prestressing loss calculation method introduced in 2005. For these components, a load factor for live load of 0.8 was maintained. Service I should be used for checking tension related to transverse analysis of concrete segmental girders. The principal tensile stress check is introduced in order to verify the adequacy of webs of segmental concrete girder bridges for longitudinal shear and torsion. Service IV—Load combination relating only to tension in prestressed concrete columns with the objective of crack control. Table 3.4.1-1—Load Combinations and Load Factors Load Combination Limit State DC DD DW EH EV ES EL PS CR SH LL IM CE BR PL LS WA WS W L FR TU TG SE Use One of These at a Time EQ BL IC CT CV Strength I (unless noted) γp 1.75 1.00 — — 1.00 0.50/1.20 γTG γSE — — — — — Strength II γp 1.35 1.00 — — 1.00 0.50/1.20 γTG γSE — — — — — Strength III γp — 1.00 1.40 — 1.00 0.50/1.20 γTG γSE — — — — — Strength IV γp — 1.00 — — 1.00 0.50/1.20 — — — — — — — Strength V γp 1.35 1.00 0.40 1.0 1.00 0.50/1.20 γTG γSE — — — — — Extreme Event I γp γEQ 1.00 — — 1.00 — — — 1.00 — — — — Extreme Event II γp 0.50 1.00 — — 1.00 — — — — 1.00 1.00 1.00 1.00 Service I 1.00 1.00 1.00 0.30 1.0 1.00 1.00/1.20 γTG γSE — — — — — Service II 1.00 1.30 1.00 — — 1.00 1.00/1.20 — — — — — — — Service III 1.00 0.80 γLL 1.00 — — 1.00 1.00/1.20 γTG γSE — — — — — Service IV 1.00 — 1.00 0.70 — 1.00 1.00/1.20 — 1.0 — — — — — Fatigue I— LL, IM & CE only — 1.50 — — — — — — — — — — — — Fatigue II— LL, IM & CE only — 0.75 — — — — — — — — — — — —

242 Table 3.4.1-4—Load Factors for Live Load for Service III Load Combination, γLL Component γLL Prestressed concrete components designed using a refined time step method to determine the time-dependant prestressing losses in conjunction with the gross section properties and without taking advantage of the elastic gain 0.8 All other prestressed concrete components 1.0

243 7.5 Fatigue 7.5.1 Proposed Revisions to Section 3 3.4—LOAD FACTORS AND COMBINATIONS • • • • Table 3.4.1-1—Load Combinations and Load Factors Load Combination Limit State DC DD DW EH EV ES EL PS CR SH LL IM CE BR PL LS WA WS WL FR TU TG SE Use One of These at a Time EQ BL IC CT CV Strength I (unless noted) γp 1.75 1.00 — — 1.00 0.50/1.20 γTG γSE — — — — — Strength II γp 1.35 1.00 — — 1.00 0.50/1.20 γTG γSE — — — — — Strength III γp — 1.00 1.4 0 — 1.00 0.50/1.20 γTG γSE — — — — — Strength IV γp — 1.00 — — 1.00 0.50/1.20 — — — — — — — Strength V γp 1.35 1.00 0.4 0 1.0 1.00 0.50/1.20 γTG γSE — — — — — Extreme Event I γp γEQ 1.00 — — 1.00 — — — 1.00 — — — — Extreme Event II γp 0.50 1.00 — — 1.00 — — — — 1.00 1.00 1.00 1.00 Service I 1.00 1.00 1.00 0.3 0 1.0 1.00 1.00/1.20 γTG γSE — — — — — Service II 1.00 1.30 1.00 — — 1.00 1.00/1.20 — — — — — — — Service III 1.00 0.80 1.00 — — 1.00 1.00/1.20 γTG γSE — — — — — Service IV 1.00 — 1.00 0.7 0 — 1.00 1.00/1.20 — 1.0 — — — — — Fatigue I— LL, IM & CE only — 1.50 2.0 — — — — — — — — — — — — Fatigue II— LL, IM & CE only — 0.75 0.80 — — — — — — — — — — — — 7.5.2 Proposed Revisions to Section 5 5.5.3 Fatigue Limit State • • • •

244 5.5.3.2—Reinforcing Bars The constant-amplitude fatigue threshold, (ΔF)TH, for straight reinforcement and welded wire reinforcement without a cross weld in the high-stress region shall be taken as: ( ) min24 20 / yTHF f f∆ = − (5.5.3.2-1) ( ) min30 25 / yTHF f f∆ = − (5.5.3.2-1) The constant-amplitude fatigue threshold, (ΔF)TH, for straight welded wire reinforcement with a cross weld in the high-stress region shall be taken as: ( ) min16 0.33THF f∆ = − (5.5.3.2-2) ( ) min20 0.41THF f∆ = − (5.5.3.2-2) where: fmin = minimum live load stress resulting from the Fatigue I load combination, combined with the more severe stress from either the permanent loads or the permanent loads, shrinkage, and creep-induced external loads; positive if tension, negative if compression (ksi) The definition of the high-stress region for application of Eqs. 5.5.3.2-1 and 5.5.3.2-2 for flexural reinforcement shall be taken as one-third of the span on each side of the section of maximum moment. C5.5.3.2 Bends in primary reinforcement should be avoided in regions of high stress range. Structural welded wire reinforcement has been increasingly used in bridge applications in recent years, especially as auxiliary reinforcement in bridge I- and box beams and as primary reinforcement in slabs. Design for shear has traditionally not included a fatigue check of the reinforcement as the member is expected to be uncracked under service conditions and the stress range in steel minimal. The stress range for steel bars has existed in previous editions. It is based on Hansen et al. (1976). The simplified form in this edition replaces the (r/h) parameter with the default value 0.3 recommended by Hansen et al. Inclusion of limits for WWR is based on recent studies by Hawkins et al. (1971, 1987) and Tadros et al. (2004). Coefficients in Eqs. 5.5.3.2-1 and 5.5.3.2-2 have been updated based on calibration reported in Kulicki et al. (2013). Since the fatigue provisions were developed based primarily on ASTM A615 steel reinforcement, their applicability to other types of reinforcement is largely unknown. Consequently, a cautionary note is added to the Commentary. 7.5.3 Proposed Revisions to Section 6 6.6.1—Fatigue 6.6.1.1—General • • • • C6.6.1.1 .

245 6.6.1.2.3—Detail Categories Components and details shall be designed to satisfy the requirements of their respective detail categories summarized in Table 6.6.1.2.3-1. Where bolt holes are depicted in Table 6.6.1.2.3-1, their fabrication shall conform to the provisions of Article 11.4.8.5 of the AASHTO LRFD Bridge Construction Specifications. Where permitted for use, unless specific information is available to the contrary, bolt holes in cross-frame, diaphragm, and lateral bracing members and their connection plates shall be assumed for design to be punched full size. Except as specified herein for fracture critical members, where the projected 75-year single lane Average Daily Truck Traffic (ADTT)SL is less than or equal to that specified in Table 6.6.1.2.3-2 for the component or detail under consideration, that component or detail should be designed for finite life using the Fatigue II load combination specified in Table 3.4.1-1. Otherwise, the component or detail shall be designed for infinite life using the Fatigue I load combination. The single-lane Average Daily Truck Traffic (ADTT)SL shall be computed as specified in Article 3.6.1.4.2. For components and details on fracture-critical members, the Fatigue I load combination specified in Table 3.4.1-1 should be used in combination with the nominal fatigue resistance for infinite life specified in Article 6.6.1.2.5. Orthotropic deck components and details shall be designed to satisfy the requirements of their respective detail categories summarized in Table 6.6.1.2.3-1 for the chosen design level shown in the table and as specified in Article 9.8.3.4. C6.6.1.2.3 Components and details susceptible to load-induced fatigue cracking have been grouped into eight categories, called detail categories, by fatigue resistance. Experience indicates that in the design process the fatigue considerations for Detail Categories A through B′ rarely, if ever, govern. Nevertheless, Detail Categories A through B′ have been included in Table 6.6.1.2.3-1 for completeness. Investigation of components and details with a fatigue resistance based on Detail Categories A through B′ may be appropriate in unusual design cases. Table 6.6.1.2.3-1 illustrates many common details found in bridge construction and identifies potential crack initiation points for each detail. In Table 6.6.1.2.3- 1, “Longitudinal” signifies that the direction of applied stress is parallel to the longitudinal axis of the detail. “Transverse” signifies that the direction of applied stress is perpendicular to the longitudinal axis of the detail. Category F for allowable shear stress range on the throat of a fillet weld has been eliminated from Table 6.6.1.2.3-1. When fillet welds are properly sized for strength considerations, Category F should not govern. Fatigue will be governed by cracking in the base metal at the weld toe and not by shear on the throat of the weld. Research on end-bolted cover plates is discussed in Wattar et al. (1985). Where the design stress range calculated using the Fatigue I load combination is less than (ΔF)TH, the detail will theoretically provide infinite life. Except for Categories E and E′, for higher traffic volumes, the design will most often be governed by the infinite life check. Table 6.6.1.2.3-2 shows for each detail category the values of (ADTT)SL above which the infinite life check governs, assuming a 75-yr design life and one stress range cycle per truck. The values in the second column of Table 6.6.1.2.3- 2 were computed as follows: ( ) ( )( )( ) 375 _ ( ) 365 75 2.5 SL TH A Year ADTT F n = ∆      (C6.6.1.2.3-1) using the values for A and (∆F)TH specified in Tables 6.6.1.2.5-1 and 6.6.1.2.5-3, respectively, a fatigue design life of 75 yr and a number of stress range cycles per truck passage, n, equal to one. These values were rounded up to the nearest five trucks per day. That is, the indicated values were determined by equating infinite life and finite life resistances with due regard to the difference in load factors used with the Fatigue I and Fatigue II load combinations. For other values of n, the values in Table 6.6.1.2.3-2 should be modified by dividing by the appropriate value of n taken from Table 6.6.1.2.5-2. For other values of the fatigue design life, the values in Table 6.6.1.2.3-2 should be modified by multiplying the values by the ratio of 75 divided by the fatigue life sought in years. Some of the values of the parameter A and the threshold (∆F)TH have been revised based on a calibration reported in Kulicki et al. (2013). The constant in the denominator of the equation to the right has been changed from 2 to 2.5.

246 The procedures for load-induced fatigue are followed for orthotropic deck design. Although the local structural stress range for certain fatigue details can be caused by distortion of the deck plate, ribs, and floorbeams, research has demonstrated that load-induced fatigue analysis produces a reliable assessment of fatigue performance. Considering the increased γLL and cycles per truck passage (n) in orthotropic decks, the 75-yr ADTTSL equivalent to infinite life (trucks per day) results in 870 for deck plate details and 4350 for all other details, based on Category C. Thus, finite life design may produce more economical designs on lower-volume roadways.

247 Table 6.6.1.2.3-1—Detail Categories for Load-Induced Fatigue Description Category Constant A (ksi3) Threshold (ΔF)TH ksi Potential Crack Initiation Point Illustrative Examples Section 1—Plain Material away from Any Welding 1.1 Base metal, except noncoated weathering steel, with rolled or cleaned surfaces. Flame-cut edges with surface roughness value of 1,000 µ-in. or less, but without re-entrant corners. A 250 × 108 24 Away from all welds or structural connections 1.2 Noncoated weathering steel base metal with rolled or cleaned surfaces designed and detailed in accordance with FHWA (1989). Flame-cut edges with surface roughness value of 1,000 µ-in. or less, but without re-entrant corners. B 120 × 108 16 Away from all welds or structural connections 1.3 Member with re-entrant corners at copes, cuts, block- outs or other geometrical discontinuities made to the requirements of AASHTO/AWS D1.5, except weld access holes. C 44 × 108 10 At any external edge 1.4 Rolled cross sections with weld access holes made to the requirements of AASHTO/AWS D1.5, Article 3.2.4. C 44 x 108 10 In the base metal at the re-entrant corner of the weld access hole 1.5 Open holes in members (Brown et al., 2007). D 22 × 108 21 × 108 7 8 In the net section originating at the side of the hole Section 2—Connected Material in Mechanically Fastened Joints 2.1 Base metal at the gross section of high-strength bolted joints designed as slip-critical connections with pretensioned high-strength bolts installed in holes drilled full size or subpunched and reamed to size—e.g., bolted flange and web splices and bolted stiffeners. (Note: see Condition 2.3 for bolt holes punched full size; see Condition 2.5 for bolted angle or tee section member connections to gusset or connection plates.) B 120 × 108 16 Through the gross section near the hole (continued on next page)

248 Table 6.6.1.2.3-1 (continued)—Detail Categories for Load-Induced Fatigue Description Category Constant A (ksi3) Threshold (ΔF)TH ksi Potential Crack Initiation Point Illustrative Examples Section 2—Connected Material in Mechanically Fastened Joints (continued) 2.2 Base metal at the net section of high-strength bolted joints designed as bearing-type connections but fabricated and installed to all requirements for slip-critical connections with pretensioned high- strength bolts installed in holes drilled full size or subpunched and reamed to size. (Note: see Condition 2.3 for bolt holes punched full size; see Condition 2.5 for bolted angle or tee section member connections to gusset or connection plates.) B 120 × 108 16 In the net section originating at the side of the hole 2.3 Base metal at the net section of all bolted connections in hot dipped galvanized members (Huhn and Valtinat, 2004); base metal at the appropriate section defined in Condition 2.1 or 2.2, as applicable, of high-strength bolted joints with pretensioned bolts installed in holes punched full size (Brown et al., 2007); and base metal at the net section of other mechanically fastened joints, except for eyebars and pin plates, e.g., joints using ASTM A307 bolts or non-pretensioned high-strength bolts. (Note: see Condition 2.5 for bolted angle or tee section member connections to gusset or connection plates). D 22 × 108 21 × 108 7 8 In the net section originating at the side of the hole or through the gross section near the hole, as applicable 2.4 Base metal at the net section of eyebar heads or pin plates (Note: for base metal in the shank of eyebars or through the gross section of pin plates, see Condition 1.1 or 1.2, as applicable.) E 11 × 108 12 × 108 4.5 In the net section originating at the side of the hole 2.5 Base metal in angle or tee section members connected to a gusset or connection plate with high-strength bolted slip-critical connections. The fatigue stress range shall be calculated on the effective net area of the member, Ae = UAg, in which U=(1- x /L) and where Ag is the gross area of the member. x is the distance from the centroid of the member to the surface of the gusset or connection plate and L is the out-to-out distance between the bolts in the connection parallel to the line of force. The effect of the moment due to the eccentricities in the connection shall be ignored in computing the stress range (McDonald and Frank, 2009). See applicable Category above See applicable Constant above See applicable Threshold above Through the gross section near the hole, or in the net section originating at the side of the hole, as applicable L c.g. x L c.g. x (continued on next page)

249 Table 6.6.1.2.3-1 (continued)—Detail Categories for Load-Induced Fatigue Description Category Constant A (ksi3) Threshold (ΔF)TH ksi Potential Crack Initiation Point Illustrative Examples 2.5 (continued) The fatigue category shall be taken as that specified for Condition 2.1. For all other types of bolted connections, replace Ag with the net area of the member, An, in computing the effective net area according to the preceding equation and use the appropriate fatigue category for that connection type specified for Condition 2.2 or 2.3, as applicable. Section 3—Welded Joints Joining Components of Built-Up Members 3.1 Base metal and weld metal in members without attachments built up of plates or shapes connected by continuous longitudinal complete joint penetration groove welds back-gouged and welded from the second side, or by continuous fillet welds parallel to the direction of applied stress. B 120 × 108 16 From surface or internal discontinuities in the weld away from the end of the weld 3.2 Base metal and weld metal in members without attachments built up of plates or shapes connected by continuous longitudinal complete joint penetration groove welds with backing bars not removed, or by continuous partial joint penetration groove welds parallel to the direction of applied stress. B′ 61 × 108 12 13 From surface or internal discontinuities in the weld, including weld attaching backing bars 3.3 Base metal and weld metal at the termination of longitudinal welds at weld access holes made to the requirements of AASHTO/AWS D1.5, Article 3.2.4 in built-up members. (Note: does not include the flange butt splice). D 22 × 108 21 × 108 7 8 From the weld termination into the web or flange 3.4 Base metal and weld metal in partial length welded cover plates connected by continuous fillet welds parallel to the direction of applied stress. B 120 × 108 16 From surface or internal discontinuities in the weld away from the end of the weld (continued on next page)

250 Table 6.6.1.2.3-1 (continued)—Detail Categories for Load-Induced Fatigue Description Category Constant A (ksi3) Threshold (Δf)TH ksi Potential Crack Initiation Point Illustrative Examples Section 3—Welded Joints Joining Components of Built-Up Members (continued) 3.5 Base metal at the termination of partial length welded cover plates having square or tapered ends that are narrower than the flange, with or without welds across the ends, or cover plates that are wider than the flange with welds across the ends: In the flange at the toe of the end weld or in the flange at the termination of the longitudinal weld or in the edge of the flange with wide cover plates Flange thickness ≤ 0.8 in. E 11 × 108 12 × 108 4.5 Flange thickness > 0.8 in. E′ 3. 9 × 108 3.5 × 108 2.6 3.1 3.6 Base metal at the termination of partial length welded cover plates with slip-critical bolted end connections satisfying the requirements of Article 6.10.12.2.3. B 120 × 108 16 In the flange at the termination of the longitudinal weld 3.7 Base metal at the termination of partial length welded cover plates that are wider than the flange and without welds across the ends. E′ 3.9 × 108 3.5 × 108 2.6 3.1 In the edge of the flange at the end of the cover plate weld Section 4—Welded Stiffener Connections 4.1 Base metal at the toe of transverse stiffener-to-flange fillet welds and transverse stiffener-to- web fillet welds. (Note: includes similar welds on bearing stiffeners and connection plates). C′ 44 × 108 12 Initiating from the geometrical discontinuity at the toe of the fillet weld extending into the base metal 4.2 Base metal and weld metal in longitudinal web or longitudinal box-flange stiffeners connected by continuous fillet welds parallel to the direction of applied stress. B 120 × 108 16 From the surface or internal discontinuities in the weld away from the end of the weld (continued on next page)

251 Table 6.6.1.2.3-1 (continued)—Detail Categories for Load-Induced Fatigue Description Category Constant A (ksi3) Threshold (Δf)TH ksi Potential Crack Initiation Point Illustrative Examples Section 4—Welded Stiffener Connections (continued) 4.3 Base metal at the termination of longitudinal stiffener-to-web or longitudinal stiffener-to-box flange welds: With the stiffener attached by fillet welds and with no transition radius provided at the termination: Stiffener thickness < 1.0 in. Stiffener thickness ≥ 1.0 in. E E′ 11 × 108 12 × 108 3.9 × 108 3.5 × 108 4.5 2.6 3.1 In the primary member at the end of the weld at the weld toe With the stiffener attached by welds and with a transition radius R provided at the termination with the weld termination ground smooth: In the primary member near the point of tangency of the radius R ≥ 24 in. 24 in. > R ≥ 6 in. 6 in. > R ≥ 2 in. 2 in. > R B C D E 120 × 108 44 × 108 22 × 108 21 × 108 11 × 108 12 × 108 16 10 7 8 4.5 Section 5—Welded Joints Transverse to the Direction of Primary Stress 5.1 Base metal and weld metal in or adjacent to complete joint penetration groove welded butt splices, with weld soundness established by NDT and with welds ground smooth and flush parallel to the direction of stress. Transitions in thickness or width shall be made on a slope no greater than 1:2.5 (see also Figure 6.13.6.2-1). From internal discontinuities in the filler metal or along the fusion boundary or at the start of the transition Fy < 100 ksi B 120 × 108 16 Fy ≥ 100 ksi B′ 61 × 10 8 12 13 5.2 Base metal and weld metal in or adjacent to complete joint penetration groove welded butt splices, with weld soundness established by NDT and with welds ground parallel to the direction of stress at transitions in width made on a radius of not less than 2 ft with the point of tangency at the end of the groove weld (see also Figure 6.13.6.2-1). B 120 × 108 16 From internal discontinuities in the filler metal or discontinuities along the fusion boundary (continued on next page)

252 Table 6.6.1.2.3-1 (continued)—Detail Categories for Load-Induced Fatigue Description Category Constant A (ksi3) Threshold (Δf)TH ksi Potential Crack Initiation Point Illustrative Examples 5.3 Base metal and weld metal in or adjacent to the toe of complete joint penetration groove welded T or corner joints, or in complete joint penetration groove welded butt splices, with or without transitions in thickness having slopes no greater than 1:2.5 when weld reinforcement is not removed. (Note: cracking in the flange of the “T” may occur due to out-of-plane bending stresses induced by the stem). C 44 × 108 10 From the surface discontinuity at the toe of the weld extending into the base metal or along the fusion boundary 5.4 Base metal and weld metal at details where loaded discontinuous plate elements are connected with a pair of fillet welds or partial joint penetration groove welds on opposite sides of the plate normal to the direction of primary stress. C as adjusted in Eq. 6.6.1.2.5-4 44 × 108 10 Initiating from the geometrical discontinuity at the toe of the weld extending into the base metal or initiating at the weld root subject to tension extending up and then out through the weld Section 6—Transversely Loaded Welded Attachments 6.1 Base metal in a longitudinally loaded component at a transversely loaded detail (e.g. a lateral connection plate) attached by a weld parallel to the direction of primary stress and incorporating a transition radius R with the weld termination ground smooth. Near point of tangency of the radius at the edge of the longitudinally loaded component or at the toe of the weld at the weld termination if not ground smooth R ≥ 24 in. B 120 × 108 16 24 in. > R ≥ 6 in. C 44 × 108 10 6 in. > R ≥ 2 in. D 22 × 108 21× 108 7 8 2 in. > R E 11 × 108 12 × 108 4.5 For any transition radius with the weld termination not ground smooth (Note: Condition 6.2, 6.3 or 6.4, as applicable, shall also be checked.) E 11 × 108 12 × 108 4.5 (continued on next page)

253 Table 6.6.1.2.3-1 (continued)—Detail Categories for Load-Induced Fatigue Description Category Constant A (ksi3) Threshold (Δf)TH ksi Potential Crack Initiation Point Illustrative Examples Section 6—Transversely Loaded Welded Attachments (continued) 6.2 Base metal in a transversely loaded detail (e.g. a lateral connection plate) attached to a longitudinally loaded component of equal thickness by a complete joint penetration groove weld parallel to the direction of primary stress and incorporating a transition radius R, with weld soundness established by NDT and with the weld termination ground smooth: With the weld reinforcement removed: R ≥ 24 in. B 120 × 108 16 Near points of tangency of the radius or in the weld or at the fusion boundary of the longitudinally loaded component or the transversely loaded attachment 24 in. > R ≥ 6 in. C 44 × 108 10 6 in. > R ≥ 2 in. D 22 × 108 21 × 108 7 8 2 in. > R E 11 × 108 12 × 108 4.5 With the weld reinforcement not removed: R ≥ 24 in. C 44 × 108 10 At the toe of the weld either along the edge of the longitudinally loaded component or the transversely loaded attachment 24 in. > R ≥ 6 in. C 44 × 108 10 6 in. > R ≥ 2 in. D 22 × 108 21 x 108 7 8 2 in. > R (Note: Condition 6.1 shall also be checked.) E 11 × 108 12 × 108 4.5 6.3 Base metal in a transversely loaded detail (e.g. a lateral connection plate) attached to a longitudinally loaded component of unequal thickness by a complete joint penetration groove weld parallel to the direction of primary stress and incorporating a weld transition radius R, with weld soundness established by NDT and with the weld termination ground smooth: At the toe of the weld along the edge of the thinner plate In the weld termination of small radius weld transitions At the toe of the weld along the edge of the thinner plate With the weld reinforcement removed: R ≥ 2 in. D 22 × 108 21 × 108 7 8 R < 2 in. E 11 × 108 12 × 108 4.5 For any weld transition radius with the weld reinforcement not removed (Note: Condition 6.1 shall also be checked.) E 11 × 108 12 × 108 4.5 (continued on next page)

254 Table 6.6.1.2.3-1 (continued)—Detail Categories for Load-Induced Fatigue Description Category Constant A (ksi3) Threshold (Δf)TH ksi Potential Crack Initiation Point Illustrative Examples Section 6—Transversely Loaded Welded Attachments (continued) 6.4 Base metal in a transversely loaded detail (e.g. a lateral connection plate) attached to a longitudinally loaded component by a fillet weld or a partial joint penetration groove weld, with the weld parallel to the direction of primary stress (Note: Condition 6.1 shall also be checked.) See Condition 5.4 Section 7—Longitudinally Loaded Welded Attachments 7.1 Base metal in a longitudinally loaded component at a detail with a length L in the direction of the primary stress and a thickness t attached by groove or fillet welds parallel or transverse to the direction of primary stress where the detail incorporates no transition radius: In the primary member at the end of the weld at the weld toe L < 2 in. C 44 × 108 10 2 in. ≤ L ≤ 12t or 4 in D 22 × 108 21 × 108 7 8 L > 12t or 4 in. t < 1.0 in. E 11 × 108 12 × 108 4.5 t ≥ 1.0 in. (Note: see Condition 7.2 for welded angle or tee section member connections to gusset or connection plates.) E′ 3.9 × 108 3.5 × 108 2.6 3.1 7.2 Base metal in angle or tee section members connected to a gusset or connection plate by longitudinal fillet welds along both sides of the connected element of the member cross-section. The fatigue stress range shall be calculated on the effective net area of the member, Ae = UAg, in which U = (1– x /L) and where Ag is the gross area of the member. x is the distance from the centroid of the member to the surface of the gusset or connection plate and L is the maximum length of the longitudinal welds. The effect of the moment due to the eccentricities in the connection shall be ignored in computing the stress range (McDonald and Frank, 2009). E 11x108 12x108 4.5 Toe of fillet welds in connected element L c.g. x L c.g. x L L (continued on next page)

255 Table 6.6.1.2.3-1 (continued)—Detail Categories for Load-Induced Fatigue Description Category Constant A (ksi3) Threshold (Δf)TH ksi Potential Crack Initiation Point Illustrative Examples Section 8—Miscellaneous 8.1 Rib to Deck Weld—One-sided 80% (70% min) penetration weld with root gap ≤ 0.02 in. prior to welding Allowable Design Level 1, 2, or 3 C 44 × 108 10 See Figure 8.2 Rib Splice (Welded)—Single groove butt weld with permanent backing bar left in place. Weld gap > rib wall thickness Allowable Design Level 1, 2, or 3 D 22 × 108 21 × 108 7 8 See Figure 8.3 Rib Splice (Bolted)—Base metal at gross section of high strength slip critical connection Allowable Design Level 1, 2, or 3 B 120 × 108 16 See Figure 8.4 Deck Plate Splice (in Plane)— Transverse or Longitudinal single groove butt splice with permanent backing bar left in place Allowable Design Level 1, 2, or 3 D 22 × 108 21 × 108 7 8 See Figure 8.5 Rib to FB Weld (Rib)—Rib wall at rib to FB weld (fillet or CJP) Allowable Design Level 1, 2, or 3 C 44 × 108 10 See Figure (continued on next page)

256 Table 6.6.1.2.3-1 (continued)—Detail Categories for Load-Induced Fatigue Description Category Constant A (ksi3) Threshold (Δf)TH ksi Potential Crack Initiation Point Illustrative Examples 8.6 Rib to FB Weld (FB Web)—FB web at rib to FB weld (fillet, PJP, or CJP) Allowable Design Level 1 or 3 C (see Note 1) 44 × 108 10 See Figure 8.7 FB Cutout—Base metal at edge with “smooth” flame cut finish as per AWS D1.5 Allowable Design Level 1 or 3 A 250 × 108 24 See Figure 8.8 Rib Wall at Cutout—Rib wall at rib to FB weld (fillet, PJP, or CJP) Allowable Design Level 1 or 3 C 44 × 108 10 See Figure 8.9 Rib to Deck Plate at FB Allowable Design Level 1 or 3 C 44 × 108 10 See Figure Note 1: Where stresses are dominated by in-plane component at fillet or PJP welds, Eq. 6.6.1.2.5-4 shall be considered. In this case, ∆f should be calculated at the mid-thickness and the extrapolation procedure as per Article 9.8.3.4.3 need not be applied. Section 9—Miscellaneous 9.1 Base metal at stud-type shear connectors attached by fillet or automatic stud welding 44 × 108 10 At the toe of the weld in the base metal (continued on next page)

257 Table 6.6.1.2.3-1 (continued)—Detail Categories for Load-Induced Fatigue Description Category Constant A (ksi3) Threshold (ΔF)TH ksi Potential Crack Initiation Point Illustrative Examples Section 9—Miscellaneous (continued) 9.2 Nonpretensioned high- strength bolts, common bolts, threaded anchor rods, and hanger rods with cut, ground, or rolled threads. Use the stress range acting on the tensile stress area due to live load plus prying action when applicable. At the root of the threads extending into the tensile stress area (Fatigue II) Finite Life (Fatigue I) Infinite Life E′ D 3.9 × 108 3.5 × 108 N/A N/A 7 8 Table 6.6.1.2.3-2—75-yr (ADTT)SL Equivalent to Infinite Life Detail Category 75-yrs (ADTT)SL Equivalent to Infinite Life (trucks per day) A 1030 B 1670 B′ 1585 C 2510 C′ 1455 D 2340 E 7515 E′ 6710 6.6.1.2.5—Fatigue Resistance • • • • C6.6.1.2.5

258 Table 6.6.1.2.5-1—Detail Category Constant, A Detail Category Constant, A times 108 (ksi3) A 250.0 B 120.0 B′ 61.0 C 44.0 C′ 44.0 D 22.0 21.0 E 12.0 E′ 3.9 3.5 M 164 (A325) Bolts in Axial Tension 17.1 M 253 (A490) Bolts in Axial Tension 31.5 Table 6.6.1.2.5-2—Cycles per Truck Passage, n Longitudinal Members Span Length >40.0 ft ≤40.0 ft Simple Span Girders 1.0 2.0 Continuous Girders 1) near interior support 1.5 2.0 2) elsewhere 1.0 2.0 Cantilever Girders 5.0 Orthotropic Deck Plate Connections Subjected to Wheel Load Cycling 5.0 Trusses 1.0 Transverse Members Spacing > 20.0 ft ≤20.0 ft 1.0 2.0 For the purpose of determining the stress range cycles per truck passage for continuous spans, a distance equal to one-tenth the span on each side of an interior support should be considered to be near the support. Values of n for longitudinal members have been revised based on the calibration reported in Kulicki et al., 2013. The number of stress range cycles per passage is taken as 5.0 for cantilever girders because this type of bridge is susceptible to large vibrations, which cause additional cycles after the truck has left the bridge (Moses et al., 1987; Schilling, 1990). Orthotropic deck details that are connected to the deck plate (e.g., the rib-to-deck weld) are subjected to cycling from direct individual wheel loads. Thus, the passage of one design truck results in five fatigue load cycles as each axle produces one load cycle. The force effect (∆f) can be conservatively taken as the worst case from the five wheels or by application of Miner’s Rule to determine the effective stress range from the group of wheels.

259 Table 6.6.1.2.5-3—Constant-Amplitude Fatigue Thresholds Detail Category Threshold (ksi) A 24.0 B 16.0 B′ 12.0 13.0 C 10.0 C′ 12.0 D 7.0 8.0 E 4.5 E′ 2.6 3.1 M 164 (A 325) Bolts in Axial Tension 31.0 M 253 (A 490) Bolts in Axial Tension 38.0