LATITUDE FOR HANDLING LONGITUDINAL FORCES. By:

Transcription

1 Small, Berry and Jenkins 1 LATITUDE FOR HANDLING LONGITUDINAL FORCES By: Gregory Small, P.Eng. UMA Engineering Ltd Kensington Rd NW Calgary, AB T2N 3S3 Phone Fax Ronald Berry, PE BNSF Railway 4515 Kansas Avenue Kansas City, KS Phone Fax William Jenkins, P.Eng. UMA Engineering Ltd Gilmore Diversion Burnaby, BC V5G 3B4 Phone Fax September 21, 2004

2 Small, Berry and Jenkins 2 ABSTRACT The authors of this paper will describe the resolution of several design and construction challenges for the first phase of the replacement of this timber approach to a shared crossing of the Fraser River near Vancouver, British Columbia. The presentation will focus on the large longitudinal forces (recommended since 1997 by clauses 2.23 of Chapter 8 and of Chapter 15 in the AREMA Manual of Railway Engineering) which required deviation from BNSF standard substructure details to address the height of the trestle and its foundation on soils of deep, loose, liquefiable saturated sands. The presentation will briefly discuss seismic issues, the interference between bracing for the new substructure and adjacent bridges, environmental concerns and the construction challenge of minimizing interference with traffic from the three adjacent railroads. (Keywords: longitudinal force, seismic, bridge foundation)

3 Small, Berry and Jenkins 3 INTRODUCTION The primary purpose of this paper is to share the thought processes and experiences of three engineers who faced a bridge design and construction project that was greatly challenged by the design longitudinal live load forces recommended in Chapter 15 and seismic provisions recommended in Chapter 9 of the 2003 AREMA Manual for Railway Engineering. PROJECT DESCRIPTION The Burlington Northern and Santa Fe Railway s (BNSF) Bridge is located in Surrey, British Columbia. Originally constructed in the early 1900 s and known as Great Northern Railway Bridge No. 77, the structure is the south approach of a three-leg wye to the Fraser River swing span crossing into New Westminster, BC (Figure 1). The other two legs are owned by the Canadian National Railway and the Southern Railway of British Columbia. The BNSF s ownership comprises 2,189 feet of principally open-deck, timber pile trestle from the south bulkhead to the beginning of Canadian National Railway ownership. The BNSF approach passes over Yale Road, the Patullo drainage channel, a Southern Railway track and an abandoned BC Electric Rail track. It also passes beneath the Patullo highway bridge and a Skytrain bridge. Except for the pile bents in and adjacent to the drainage channel, the remaining bents were posted in The vertical alignment of the structure is a 0.64% uphill grade to the north, and the horizontal alignment is a compound curve of about 5 30.

4 Small, Berry and Jenkins 4 Train operation on the bridge is governed by an interlocking, locally controlled by the Fraser River bridge operator. The permanent speed restriction on the structure is 15 mph for passenger trains and 10 mph for freight trains. The signal at the north end of BNSF s leg may stop trains short of the main river bridge to protect bridge openings or train traffic on the other legs. In 2003, BNSF operated approximately 16 million gross tons across this bridge. Recent inspections of the bridge revealed evidence that the framed timber bents were tilting down-grade. UMA Engineering, Ltd. was retained in 2002 to perform a bridge situation survey. After reviewing the survey and other inspection data, BNSF expanded UMA s scope of services to include subsurface soil investigations, structural design, contract documents, permitting and field inspection of the reconstruction of 225 feet of the north end of the approach bridge. DESIGN Design Considerations BNSF required the new bridge to be designed to meet the requirements of the AREMA Manual for Railway Engineering (Manual) chapters 8, 9 and 15 for a Cooper E-80 design loading and impact for ballasted decks. The replacement structure was to be ballasted deck and of permanent (steel/concrete) construction. BNSF standard details were to be used as much as possible. The spacing of the existing timber bents was approximately 14 feet center to center, requiring the replacement structure to use span lengths in multiples of approximately 14 feet.

5 Small, Berry and Jenkins 5 Because the Fraser River crossing operates a swing span to pass marine traffic, and because it is shared between three railroads, a large proportion of northbound BNSF trains must stop with the head-end locomotives positioned on the section of new bridge to await authority to proceed across the river spans. As a result, the BNSF approach trestle sees large braking and tractive forces for a high proportion of all trains. This, along with the grade of the bridge, low operating speed, and evidence of the existing timber bents leaning, mandated that the structure be designed to carry the full magnitude of longitudinal live load force recommended in section of Chapter 15 Steel Structures of the 2003 AREMA Manual (1). Vancouver is one of the highest risk seismic zones in Canada, so the new bridge also had to be designed to withstand significant seismic activity. The section of bridge being replaced is flanked on both ends by pile and frame timber trestles. It therefore has greater mass and is considerably stiffer so that it will tend to attract longitudinal loadings from the adjacent structures, in addition to what is applied within its length. The bridge section is approximately 35 feet high from top of tie to groundline, and does not have massive structures such as abutments or piers near the level of the track to anchor horizontal forces. It therefore requires bracing to carry the longitudinal and lateral forces 35 feet below the deck to ground level. The soils beneath the trestle consist of loose, seismically liquefiable, sands and silts to a depth of 25 feet followed by a layer of compact silty sands to a depth of more than 120 feet. For lateral

6 Small, Berry and Jenkins 6 loadings, the geotechnical report recommended a lateral modulus of subgrade reaction of 100 kcf for the first 25 feet, 700 kcf to a depth of 65 feet, and 1000 kcf below this. Selection of the Bridge Structure Due to the approximate 14 foot spacing of the existing timber bents, the most economical steel/concrete bridge was determined to be eight identical spans supported on new bents spaced at 28-2 centers. (Figure 2). Prestressed concrete box T-girders were selected for the main spans because the narrow stem permitted better access for bearing connection to transfer horizontal forces and because of BNSF s negotiated volume pricing with its prestressed concrete supplier for girders of this design. The chosen substructure consisted of nine bents at 28-2 centers, driven through the open deck of the bridge, and alternately braced into four supporting towers (Figure 3). Because the piling needed to be driven through the deck between the existing stringers, a five pile configuration was designed. The geotechnical conditions required the piling to penetrate approximately 80 feet into the ground to carry the vertical loads past the seismically liquefiable layers, resulting in piles that were approximately 110 feet. long from cutoff to tip. Steel H-piles were chosen for ease of driving, splicing, and bracing into towers to carry the longitudinal and lateral loads Steel cap beams were selected for ease of erection on the 30 ft. high bents. In the experience of the principal design engineer, cap beams are very sensitive to proper fatigue detailing, so the bottom flange was stripped full length on one side to meet the face of the piling rather than

7 Small, Berry and Jenkins 7 simply coping slots to pass the piles. Furthermore, connection between the cap and the piles was bolted, and no tack or other welding of any sort was permitted during the installation. Elastomeric bearings were selected due to their low cost and successful performance under short concrete spans at BNSF. Longitudinal Forces The longitudinal live load forces recommended by AREMA (and predecessor AREA) have changed significantly over the years. A more complete history is given in Section of the Chapter 15 commentary (2), but a very brief outline of the recent evolution to current requirements is given below to explain the change in design philosophy that was developed for this bridge. Prior to 1997, Chapter 15 of AREA (3) specified that the section of bridge being designed must carry a longitudinal force equal to 15% of the weight of the Cooper consist, (15% being the assumed coefficient of kinetic friction for steel wheel on steel rail and therefore the maximum longitudinal force that could be transferred from a train going into an emergency brake application). Although not applicable to the BNSF portion of the Fraser River Bridge (due to the presence of rail joints at the swing span), if the rail was continuous from embankment to embankment, the pre 1997 AREA Chapter 15 permitted longitudinal load to be further reduced by multiplying it by the lesser of 0.8 or L/1200m, where L = the total length of the bridge.

8 Small, Berry and Jenkins 8 The advent of high adhesion diesel-electric locomotive technology has greatly increased the amount of tractive friction that locomotives can achieve, both in pulling and dynamic braking, so the AAR conducted tests in 1996 which proved the need for new requirements that became part of the 1997 edition of AREA. (4), (5) Since 1997, AREA/AREMA has recommended bridges to be designed to carry the greater of the braking force or locomotive traction force, as specified by the following two formulae (1): Longitudinal braking force (kips) = L L(ft.) (Equation 1) Longitudinal Traction force (kips) = 25 L L(ft.) (Equation 2) In comparison to the pre-1997 use of 15% of the Cooper Live Load, Equation 1 increases loading on spans that are shorter than the Cooper axle consist, but gives essentially the same force for longer spans. Equation 2 for the tractive force from the locomotives, governs for all bridge lengths between 4 and 355 ft. Figure 4 graphs the longitudinal forces specified in AREMA/AREA Chapter 15 section before and after 1997, and illustrates the substantial increase in design longitudinal forces on shorter sections of bridge. Figure 5 shows the longitudinal forces per foot of bridge before and after 1997, clearly showing how much the design unit longitudinal forces have increased for shorter bridge sections. Considering a 28 ft span, the longitudinal force has increased by a factor of 2.5.

9 Small, Berry and Jenkins 9 Figure 5 also points out an important design strategy for carrying longitudinal forces. Notice that the loading per foot of bridge length under the post 1997 AREMA requirements drops off dramatically as the length of bridge increases. If we design the structure to share the longitudinal braking force equally across as much bridge as possible (in this case, the entire 225 foot length), then the longitudinal braking force can be reduced to 1.67 kips/ft. In the experience of the principal design engineer, similar structures were traditionally designed in the past so that each tower supported the longitudinal forces for two spans (for a total length of about 56 feet in this case resulting in a design load of 3.34 kips/ft longitudinal force according to current AREMA requirements, or double the force than if we can tie together all four towers). Under the pre-1997 recommendations, load sharing across 4 towers only provided a 20% reduction in the longitudinal force per tower and was seldom done. Preliminary analysis of the ability of the piling to carry the longitudinal forces into the soil had shown that this force would govern the sizing of piles, and that tying the substructure together to equally share the longitudinal loads was needed to keep the pile sizes within practical limits and to avoid the cost of installing double bents. Note that the costs for installing double bents vs. connecting the substructure into a single contiguous structure were not compared since the consensus opinion of the design team was that the latter was obviously much less expensive due to required pile lengths. Seismic Design Loads Due to the height, foundation soils and seismic potential at this location, a sub-consultant engineering expert in Seismic design, Dr. Mahmoud Rezai of the University of British Columbia,

10 Small, Berry and Jenkins 10 was retained to assist with prediction of the seismic forces that could be seen by the structure. Note that the current seismic design requirements contained in Chapter 9 of AREMA (6) do not require seismic analysis to include the presence of a train on the bridge. Nonetheless, a greater than usual proportion of trains will be stopped on the bridge awaiting authority to proceed across the Fraser River crossing, so there was a foreseeable probability that there could be a train on the bridge during an earthquake. It was also recognized that a train is only loosely connected to the bridge structure, and it would behave as a damped, loosely sprung mass. After discussion with Dr. Rezai, the design team instructed him to model the presence of a train on the bridge as a rigidly connected mass equivalent to the heaviest 56 feet of a Cooper E-20 consist (25% of the E-80 design loading). Dr. Rezai s analysis (Figure 6) gave longitudinal seismic loadings of just under 100 kips per braced tower, approximately the same magnitude as the longitudinal live load force. Lateral Seismic loading was just under 90 kips per braced tower, which were greater than the centrifugal and wind loads. Longitudinal Struts Because the span bearings were elastomeric, there was no ability to use the spans to rigidly connect the four towers together to ensure that longitudinal forces on the 225 foot section of bridge would be equally shared. The solution was to install longitudinal struts along the tops of the towers to connect the nine bents together, as shown in Figure 3. It was decided also to strut the bents at the groundline to allow load sharing at that level as much as possible. This was done by using the same bottom strut in the towers to span the space between towers, also shown in

11 Small, Berry and Jenkins 11 Figure 3. Note that because of the drainage canal between bents 6 and 7, a groundline strut could not be placed across that opening. Thermal Considerations The thermal stresses and movements of the strutted length of 255 feet were not anticipated to cause any problems with this structure for the following reasons: 1. The spans themselves have fixed bearings on one end and expansion bearings on the other and will be expanding/contracting with the substructure (although there will be delayed reactions due to the higher thermal intertia of the concrete boxes). 2. Groundline thermal movements have never been an issue in the experience of the design team. Thermal expansion of the struts will move the piles until they react sufficiently to hold the strut s axial forces. Thus the actual physical expansion and contraction will be less than what would occur if the strut were unrestrained. Since groundline strutting was interrupted by the drainage canal, the maximum continuous length of steel at groundline was only 141 feet. 3. The bridge is on a 5.65 degree curve, so thermal movements at the tops of the towers, if unrestrained, will smoothly shift the bridge superstructure to a different curvature.

12 Small, Berry and Jenkins Temperature variation in Vancouver is approximately 60% of the range experienced by bridges in other areas of North America that are inland from the moderating effect of the Ocean. As expected, there have been no reported problems or discernable observations of thermal expansion/contraction of the bridge since completion of the construction in November Piling Selection It was not difficult to obtain the required axial capacities of the bent piling, including uplift of some piles during seismic events. The greatest challenge was carrying the horizontal forces (longitudinal and lateral) into the ground. It was the horizontal force transfer criteria which governed the selection of the size and quantity of piling. The methodology of Broms (1964) was used to estimate pile capacity to transmit horizontal loadings into the ground (7). The Broms formulation predicts pile deflection at ground level to resist an applied horizontal load, given the geometry of the loading, the stiffness properties/length of the pile and the modulus of subgrade reaction for the soils. Broms suggests a design maximum ground level movement of ¼ inch which became our criteria to estimate horizontal pile capacities. The chart produced by Broms (Figure 7) gives a dimensionless relationship between the horizontal pile deflection and the effective length of piling in the ground. Only the two extreme cases of a free headed and a fixed headed pile are addressed in his formulation. The substructure

13 Small, Berry and Jenkins 13 for the BNSF bridge is somewhere between these two extremes due to the stiffness provided by the braced towers. Therefore, the horizontal pile capacities for the two load cases were examined, and compared to the needed capacity for each braced tower to carry a horizontal force of 100 kips in the longitudinal direction (5 piles in 2 bents must carry 10 kips each). Several calculations using various sizes of piles were tried before selecting HP14x117 piles. The Broms method predicted that if this was a free headed pile, with the horizontal load applied at the level of the bottom struts (18 above the ground), that each pile had a resistance of 8.5 kips. If the head of the pile was completely fixed from rotating at the ground level, each pile would have a horizontal resistance of 31 kips. Although there was no interpolation method suggested by Broms, the design team judged that the partial fixity against pile rotation provided by the tower bracing would give a horizontal capacity somewhere between these two values, and was certainly greater than the required 10 kips. Since the tower bracing only connected to the outside piles in each bent, it was necessary to design a system to rigidly distribute the longitudinal forces to all piles across the bent, forcing them to move together to share and resist the forces. The chosen solution was a pair of C12x30 channels welded on either side of each bent at the level of the bottom longitudinal struts (see Figures 8 and 9). Batten plates were added to create a box structure that was both flexurally and torsionally rigid. The stiffness of this distribution beam was calculated to limit the relative deflection of piles within the bent to approximately 3/16 inches when a 25 kip horizontal load was applied to each of the outside piles by the bracing system. In the lateral direction, the batter of the outside piles was sufficient to carry the design lateral loadings.

14 Small, Berry and Jenkins 14 CONSTRUCTION Because the existing structure needed to remain operational during construction, and it spanned a fisheries sensitive watercourse and there was limited working area, construction of the bridge presented several interesting challenges. Regular daily Amtrak service between Seattle and Vancouver had to be accommodated as well as ongoing traffic on the BNSF track and the adjacent two railroads. Construction work was carried out by Fraser River Pile & Dredge Ltd., an experienced local general contractor. They constructed the new steel bridge substructure between and below the existing timber trestle using train free window periods no greater than six hours (between Amtrak service) and frequently less (due to adjacent CN Rail or Southern Rail traffic). All the bents, cap beams, bracing and struts for the substructure needed to be in place before any of the new prestressed concrete girders, ballast and rail could be installed. The primary challenges were: 1) Constructing the new substructure while accommodating the existing structure and full rail operations. This necessitated stringent work site safety procedures for active tracks. As well, the existing timber bents required rebracing to obtain working room and interim bracing of the new steel bents was necessary to maintain the structural integrity of the existing timber bridge during construction (Figure 10).

15 Small, Berry and Jenkins 15 2) Constructing the structural design elements while still accommodating adjacent facilities, site constraints and other railroad operations. Where the new steel bridge structure connected to CN Rail s timber structure, alignment adjustments were required for the nearest bent and the supporting struts (Figure 11). 3) Maintaining precise accuracy of the piles and cap beam locations while constructing at height with limited working room. These factors affected construction productivity but ultimately the structure was completed as desired (Figure 12). CONCLUSIONS As technology increases the tractive and braking force potential of modern freight locomotives, railroad bridges will begin to experience greater horizontal forces. Owners and bridge designers should be aware of these increasing forces as well as the site and operational conditions that increase the likelihood of the maximum design forces occurring. At BNSF Bridge 140.8, the strategy of tying the four towers of the bridge together to carry the longitudinal forces resulted in a more economical substructure than if designed according to traditional methods wherein each tower with two spans is designed as a separate unit to resist longitudinal forces. To accomplish this objective, the bracing system for the substructure included strutting of the tops of all bents, strutting of the groundline, and the construction of a flexurally and torsionally stiff distribution beam at the base of each bent.

16 Small, Berry and Jenkins 16 REFERENCES (1) AREMA chapter 15, Longitudinal Forces (2003) (2) AREMA chapter 15 commentary, Longitudinal Forces (2003) (3) AREA chapter 15, Longitudinal Forces (1996) (4) Otter, Duane E., Sweeney, Robert A.B, and Dick, Stephen, M., Development of Design guidelines for Longitudinal Forces in Bridges, TTCI Technology Digest , AAR, August 2000 (5) Uppal, A.Shakoor, Otter, Duane E., Sweeney, Robert A.B, Joy, Richard B., LoPresti, Joseph A., Maal, Daina Olivia, and Doe, Brian E., Longitudinal Forces in Bridges Due to Revenue Service Traffic, TTCI Technology Digest , AAR, August 2000 (6) AREMA chapter 9, Seismic Design for Railway Structures (2002) (7) Broms, Bengt B., Lateral Resistance of Piles in Cohesionless Soils, ASCE J. SMFE, V90, SM 3, May 1964

17 Small, Berry and Jenkins 17 LIST OF FIGURES Figure 1 Aerial View of the Fraser River Bridge Figure 2 Typical Cross section of Superstructure Figure 3 Bridge Profile Figure 4 - Comparison of Pre and Post 1997 Longitudinal Force Requirements Figure 5 - Comparison of Pre and Post 1997 Unit Longitudinal Forces Figure 6 Seismic Analysis of one Tower and Two Spans Figure 7 Broms Dimensionless Chart of Groundline Deflection vs. Effective Pile Length Figure 8 Bent Bracing Figure 9 Bottom Distribution Beam Figure 10 Construction Work Figure 11 Alignment Adjustments Figure 12 Maintaining Piles and Cap Beam Locations

18 Small, Berry and Jenkins 18 Figure 1 Aerial View of the Fraser River Bridge

19 Small, Berry and Jenkins 19 Figure 2 Typical Cross section of Superstructure Figure 3 Bridge Profile

20 Small, Berry and Jenkins 20! " #! " # " " # $ " % &'( ')! " # " " # $ " % &'( ')*! " # " " # $ " % &'( ') Figure 4 - Comparison of Pre and Post 1997 Longitudinal Force Requirements %! " # & $ )! " # " " # $ " % &'( ')! " # " " # $ " % &'( ')*! " # " " # $ " % &'( ') Figure 5 - Comparison of Pre and Post 1997 Unit Longitudinal Forces

21 Small, Berry and Jenkins 21 Figure 6 Seismic Analysis of one Tower and Two Spans

22 Small, Berry and Jenkins 22 Figure 7 Broms (7 )Dimensionless Chart of Groundline Deflection vs. Effective Pile Length

23 Small, Berry and Jenkins 23 Figure 8 Bent Bracing Figure 9 Bottom Distribution Beam

24 Small, Berry and Jenkins 24 Working Close Driving Piles Figure 10 - Construction Work

25 Small, Berry and Jenkins 25 Cap Beam Bottom Distribution Beam Figure 11 Alignment Adjustments

26 26 Small, Berry and Jenkins Bents Struts and Bracing Figure 12 Maintaining Piles and Cap Beam Locations