EVALUATION OF THE LOAD CAPACITY OF A REHABILITATED STEEL ARCH RAILWAY BRIDGE

Transcription

1 EVALUATION OF THE LOAD CAPACITY OF A REHABILITATED STEEL ARCH RAILWAY BRIDGE J.F. Unsworth, P.Eng. Manager, Structures Planning and Design, Canadian Pacific Railway, Suite th Ave. SW, Calgary, Alberta, Canada Tel Fax

2 EVALUATION OF THE LOAD CAPACITY OF A REHABILITATED STEEL ARCH RAILWAY BRIDGE J.F. Unsworth 1 1 Manager, Structures Planning and Design, Canadian Pacific Railway, Calgary, Canada ABSTRACT The load carrying capacity of a 336 foot span steel arch bridge carrying heavy freight railway traffic, originally constructed in 1893 and extensively strengthened in 1929, was recently evaluated through detailed visual inspection, non-destructive testing and three-dimensional finite element structural analysis. This paper reviews the load capacity evaluation of the bridge and the results of the evaluation in light of present railway traffic loads on the structure. The evaluation shows that the bridge is adequate to safely carry modern railway traffic. Also, based on the load rating, a simple procedure for the evaluation of infrequent very heavy loads can be established. INTRODUCTION Canadian Pacific Railway s Stoney Creek bridge is a single track steel deck arch railway bridge spanning a 300 foot deep canyon about 200 miles west of Calgary, Alberta, near the summit of Canada s Selkirk Mountain Range (Fig. 1). The bridge consists of two parallel three-hinged steel pin-connected truss arches spanning 336 feet with a rise of 100 feet on each side of the railway track. The parallel truss arches are at a batter of 1:10 (H:V), are 48 feet apart at the base and taper to 28 feet at the crown. This arrangement of the arches creates an elegance of form as well as providing considerable lateral stiffness. The track is supported by deck plate girder spans on spandrel bents. The single railway track is tangent across the bridge, with the exception of a horizontal spiral curve commencing at the west quarter of the arch. The spiral curve provides a transition to an 11 degree simple curve starting about 50 feet off the west end of the bridge. The bridge is on a grade of 2% ascending westward.

3 The Original Stoney Creek Bridge (1885) The original railway bridge at this location was a four span wooden Howe truss supported on timber tower piers constructed in 1885 and designed by prominent American civil engineer C. C. Schneider. With its 230-foot high central tower, the bridge was reputed to be the highest railway bridge in the world (Lavallée 1974). However, Gustave Eiffel's Garabit Viaduct, constructed in 1884 of wrought iron, rises 400 feet above the Truyère River in France, disputing this claim. Nevertheless, it may have been the highest timber bridge constructed. The introduction of heavier locomotives, considerable maintenance requirements associated with large timber bridges and risk of destruction by fire precipitated its replacement with a steel arch bridge after only eight years of service. The Steel Arch Bridge (1893) The steel arch bridge was constructed in 1893 on falsework within the deep canyon. The bridge consisted of single three-hinged steel pin-connected truss arches spanning 336 feet with a rise of 100 feet on each side of the railway track. The track over the arch was supported by seven deck lattice girder spans on spandrel bent columns. The spandrel bent columns connected to every third panel point from each end of the truss arch and deliver reactions to the arch at six points (Fig. 2). In order to create a statically determinate three-hinged structure, the main hinge pins are located at the crown and ends of the arch at the lower chord only with a sliding pin arrangement at the crown of the top chord. This arrangement precludes axial stress in the top chord members in the center panels but makes it necessary that the bottom chord of the truss arch carry most of the live load. However, the three-hinged design allows for a better distribution of forces between top and bottom chords of the truss arch than a two-hinged arch arrangement. The forces in the top chord of the Stoney Creek arch are about 20% of that in the bottom chord at the end supports. In contrast, the forces in the top chord of the Hell Gate arch bridge in New York City (a two-hinged steel truss arch of similar design to Stoney Creek) are about 6% of that of the lower chord at the end supports (Ammann 1918). The bridge was designed for a live load consisting of two consolidation locomotives with axle

4 loads of 30 kips, 48 kips and 42.5 kips (196 ton nine axle locomotives) and a uniform train load of 3000 lb/ft without a dynamic load allowance (Fig. 2). Lateral wind forces of 360 lb/ft applied 7 feet above the rail on the moving train and 30 lb/ft 2 on the bridge were used for lateral bracing design. Lateral wind forces for the unloaded bridge condition were taken as 50 lb/ft 2. The wind pressures were applied to twice the elevation surface area of the bridge. Arch truss member forces due to wind were superimposed on member forces due to dead and live load only when the member forces due to wind were in excess of 25% those due to dead and live load. Longitudinal forces due to tractive effort and braking were calculated as 20% of the vertical live load. The steel was specified to have an elastic limit of 33 ksi (Motley 1930). Allowable tension and compression design stresses for the truss arch member were specified sufficiently low to account for dynamic load amplification and preclude elastic instability. Typical specification for truss chord allowable tension was 12 ksi and for allowable compression L R 2 ksi ; where L= length of the compression member (in.) and R= least radius of gyration (in.). After construction, vertical arch deflections of less than ¾ in. at the crown were reported under usual train loads of 1893 (Engineering News 1894). The Reinforced Steel Arch Bridge (1929) Further increases in locomotive weights required that the railway either replace or rehabilitate the Stoney Creek bridge. Initially, it was decided to replace the 1893 steel arch with a new 311 ft cantilever deck truss with 111 ft flanking anchor spans adjacent to the existing steel arch. However, foundation conditions in the canyon curtailed the economical construction of a new bridge and rehabilitation of the existing 1893 construction was undertaken in The rehabilitation consisted of the erection of two additional truss arches 5 feet to the outside of the existing truss arches. The wider overall bridge geometry provides a greater lateral stiffness to resist lateral loads due to wind, centrifugal force and live load. Longitudinal forces due to traction and braking are transmitted to the arches through new latticed struts between spandrel bents. The new trussed arches were designed to be identical in geometry with member sections and

5 materials of close similarity to the existing 1983 arches (Motley 1930). Furthermore, the arches were tied together with pinned diaphragms to provide for a uniform stiffness and distribution of live load to each pair of arches. Equalizing diaphragms were designed to connect between the top chords of the parallel arches at each panel point in the end quarter spans and at every second panel point in the central portion of the arches. This arrangement transfers spandrel bent column loads to each arch at thirteen locations (Fig. 3). However, due to construction constraints and the shortening of the track spans supported by the spandrel bents, the original end spandrel bent columns were not replaced and remain supported by the original 1893 arches only. The spandrel bent columns are pinned at the top and have disc-type bearings at the base where they connect to the panel point equalizing diaphragms. Equalization of loading also required burdening the new arches with dead load at each panel point corresponding to the dead load existing on the 1893 arches before connecting the trusses. This necessitated construction of the new arches at higher elevations than the existing arches in accordance with the calculated dead load deflection of 5/8 in. at the centre of the arches (Motley 1930). Equalized dead load deflections were confirmed by field measurements at the time of construction. After construction, deflections under a static test live load consisting of four locomotives (ranging in weight from 230 to 376 tons) with a total weight of 1100 tons showed a uniform deflection of 1-5/8 in. at the arch centres. It is interesting to note that under the usual inservice train loading conditions of 1929, the original 1893 arch deflected 2-3/4 in. at the centre pin (Motley 1930). An analysis for Cooper s E-60 design live load (Fig. 3) illustrated that some members of the original arches required reinforcement. It is generally considered that the rehabilitation improved the bridge aesthetics as well as structural behaviour.

6 EVALUATION OF THE REHABILITATED BRIDGE UNDER MODERN TRAIN LOADS Bulk commodity unit train traffic introduced in the early 1970 s created train axle loads of almost 70 kips (200 ton six axle locomotives and 110 ton four axle cars). An evaluation was made in 1970 based on parameters of the original design and in accordance with American Railway Engineering Association recommended methods of the time. These methods are similar to the "maximum" load rating methodology outlined in the American Railway Engineering and Maintenance-of-Way Association Manual of Recommended Practice (AREMA 1998) current at the time of the investigation (Conway 2001). The 1970 load rating, using a yield stress of 30 ksi, exhibited that all elements of the structure were adequate for train loads less than Cooper s E-104 live load at 60 mph. The governing demand to capacity (D/C) of arch members was 0.77 for member TC2. Regular unit freight train traffic with car axle loads of 72 kips and business pressures to increase car commodity volumes that create 80 kip axle loads, precipitated a further investigation of the structure using modern methods of analysis and strength rating in 1998 and Furthermore, recent inspections revealing vertical elongation of the bottom pin holes at the 1893, and to a lesser extent in the 1929, truss arch vertical end members, indicated a need to determine the behaviour of the bridge in greater detail. The 1999 strength evaluation consisted of three-dimensional modelling the structure (Fig. 4) with linearly elastic frame elements (Buckland and Taylor 2000). Connections between the 1893 and 1929 trusses were modelled with appropriate pinned end connections. It was recognized that the deformations associated with the elongated pin holes at the end verticals of the arch generate geometrically non-linear structural behaviour. However, it was considered appropriate to proceed with a linear elastic analysis based on deformations producing only slight non-linearity that would not invalidate the linear elastic analysis for purposes of a practical strength rating (Cook 1981). The evaluation investigates the load distribution, strength and behaviour of the bridge under a Cooper s E-80 loading. Cooper s E-80 loading is identical to that shown in Figure 3 with axle loads increased proportionally by 33%. The load rating was carried out in accordance with AREMA methods for "normal" and

7 "maximum" ratings. "Normal" ratings are based on design allowable stresses and "maximum" ratings on allowable stresses of up to 80% of specified material yield stress. At Canadian Pacific Railway "normal" ratings are generally used to assess existing structures for safe passage of traffic over the expected service life of the bridge and "maximum" ratings are used to assess the effects of infrequent loads. "Maximum" ratings may also be used to assess structures for usual railway traffic where traffic volumes are such that the remaining useful life of the structure is not substantially affected (AREMA 1998). The yield stress of the steel used in the determination of member capacity was 30 ksi, but reduced by 15% for the 1893 truss due to some uncertainty regarding the actual material properties of the 1893 steel (Motley 1930). The demand on a member or pin, from a Cooper s E- 80 loading at a train speed of 60 mph, was calculated by the three-dimensional elastic frame analysis. Deficient members were reviewed for demand with reduced centrifugal and dynamic (impact) loads at reduced speeds to determine the permissible maximum train speeds for the Cooper s E-80 loading. If train speeds cannot be reduced to attain a D/C 1.0 then the maximum equivalent Cooper s loading for the D/C=1.0 condition is developed for a 60 mph train speed. SUMMARY OF RESULTS OF THE EVALUATION The three-dimensional analysis revealed that an unequal distribution of load is being transferred to each arch. Furthermore, at the west end of the bridge the demand requirements on the arches is considerably greater due to the centrifugal effects from track curvature. The centrifugal force, F c (kips), is; F C 2 6.7PV 100R = (1) where P =axle load (kips); V = train speed (mph); and R = radius of track curvature (ft), which varies from R = at the beginning of the curve at the west quarter point of the truss (P4) to R= 695 ft at the west end of the bridge. In accordance with AREMA specifications this horizontal force is applied at each axle location at a point 6 feet above the rail. All elements of the bridge are of sufficient strength for Cooper s E-80 loading at 60 mph at "maximum" rating member capacity, with exception of the west 26 ft deck plate girder span and

8 the main pin for the north 1929 truss arch at the west end. The results of the analysis showed D/C values of 0.85 and 0.74 for "maximum" rating of the 1893 and 1929 arch members TC2, respectively. Member TC3, with a D/C=0.89, governed the "maximum" strength rating for truss arch members. The moderate differences between the 1970 and 1999 strength evaluations are primarily due to the assumption of equal distribution of live load to each arch and the omission of centrifugal effects in the 1970 assessment. However, the "normal" load rating indicated that, for a Cooper s E-80 loading, many members of the bridge are of deficient capacity, as shown in Table 1. "Normal" rating results show that all truss arch diagonals, truss member end pins and arch main pins (all pins were rated for the governing effects due to coexisting member forces) are of sufficient capacity for Cooper s E-80 loading at 60 mph. However, many other members of the truss arches are of deficient capacity under Cooper s E-80 loading at 60 mph. The two spandrel bent transverse beams that support a 21 ft. and 42 ft. deck plate girder span are also deficient at "normal" rating stresses. In addition, the four deck plate girders at the west end of the bridge are deficient at "normal" rating stresses. All other spandrel frame members, transverse beams and deck plate girders are sufficient at "normal" rating stresses. DISCUSSION OF RESULTS "Maximum" Rating Results The 26 ft. deck plate girder span at the west end of the bridge has a D/C=1.37 at "maximum" rating stresses. The governing criteria is flexural stress at mid-span. For loads equivalent to Cooper s E-80, the demand can be reduced, so that D/C 1.0, through a reduction in centrifugal force and impact by limiting train speeds to 30 mph. This information may be used to establish speed restrictions or impasses to special or infrequent heavy loads by establishing the equivalent Cooper s E load, based on mid-span flexure, for the load being considered. The main pin of the north 1929 truss arch at the west end has a D/C=1.03 for "maximum" rating based on bearing stresses between the main pin and pin plates. The AREMA Manual allows that bearing stresses on pins may be disregarded for bridge rating unless there is visible deformation of contact parts. Visible deformations have been reported at these pins. As proper

9 functioning of these pins at the hinged skewbacks is of primary importance to the behaviour of the arch, and because AREMA does not provide for explicit "maximum" pin bearing rating stresses, allowable stresses corresponding to a "normal" rating were used to assess the capacity. Therefore, the D/C values reported for the main arch pins at the base can be considered conservative and disregarded with respect to the evaluation of infrequent heavy loads of the bridge. "Normal" Rating Results The situation is a more complex with respect to "normal" rating. Normal rating results indicate that either strengthening of deficient members or load reductions are required. A combination of selective member strengthening in conjunction with application of a maximum train speed that does not compromise the efficiency of train operations is a cost effective means of ensuring safe load carrying capacity of the bridge. Effective reductions in live load can be reached through speed restriction developed in accordance with the formula given by AREMA for centrifugal force (Eq. 1) and reduction of dynamic effects, I R, (Eq. 2). Eq. 2 is applied to only the vertical effects components of the AREMA impact equation. 2 ( ) 0.80 IR = 1 60 V (2) 2500 Table 2 shows deficient members for the "normal" rating and speed reductions required to safely pass Cooper's E-80 loads. In some cases, even a static application of the Cooper's E-80 load creates a demand that exceeds member capacity. Due to track grade and curvature at the bridge, current train speeds are limited to 25 mph. Table 3 shows the maximum Cooper's E axle loading for members that remain deficient for Cooper's E-80 loading below 25 mph. The pins at the base of the truss were shown to have D/C values approaching 1. Observations of pin hole elongation at the base of the truss end verticals would appear to indicate that, prior to the 1929 rehabilitation, elongations were likely occurring; and may have continued after the strengthening due to the end spandrel bent being supported directly over the 1893 arches. Based on the "normal" rating results, train loads should be limited to E-59 as governed by member TC3 at the west end of the 1893 trusses.

10 Fatigue Consideration In 1988 a new track was constructed lower down the canyon and the Stoney Creek bridge loading regime became primarily that of empty eastbound traffic. A life cycle fatigue analysis (CPR 1992) using modern methods of stress range prediction and damage accumulation rules (Dick & McCabe 1991) was conducted. The analyses indicated that, until regular unit freight train operations commenced in 1974, very few stress range cycles with magnitude greater than the tension member constant amplitude fatigue stress thresholds were likely developed. However, even though fatigue damage accumulation between 1974 and 1988 was considerable for the short span girders, it did not approach the theoretical fatigue life. Since 1988, traffic volume and loading reductions (primarily empty unit trains) on the bridge have reduced the frequency of equivalent constant amplitude stress range magnitudes (computed as the root mean cube stress range using Rainflow cycle counting technique) above the fatigue stress limit for these riveted members; and a further a study of remaining fatigue life was not deemed necessary. CONCLUSIONS AND RECOMMENDATIONS General Non-uniform distribution of live load to the trusses, in conjunction with lateral forces due to centrifugal effects at the west end of the structure, establish the governing strength rating of the bridge. The existence of relatively high bearing stresses at the main pins at the end of the arch and the observation of pin hole elongation and material bearing failure, indicates that periodic nondestructive testing and further investigation (geometric non-linear analysis) of local conditions and effect on the overall structural bevavior should be considered. "Maximum" Rating Results The passage of infrequent heavy loads across the bridge is governed by mid-span flexure of the 26 ft. deck plate girder span at the west end of the bridge. Therefore, a simple determination of the safety of infrequent heavy loads across the bridge is available. "Normal" Rating Results Several bridge components (Fig. 5) were found to be deficient at "normal" rating stresses with

11 respect to the demand created by Cooper s E 80 load at a track speed of 25 mph. However, with the present lower traffic volumes and train weights, the remaining useful life of the structure is not substantially affected and the immediate consideration of load carrying capacity for usual traffic may be based on "maximum" ratings. This indicates that the bridge is currently safe to carry train loads equivalent to Cooper's E-80 at 25 mph. However, in order to ensure the capacity of the structure to safely sustain present and future train loads, the deficient members in Table 3 may require strengthening. In order to effectively plan an appropriate maintenance program, an analysis of the equivalent Cooper's E rating for present train configurations for the deficient members was performed. The results of that analysis indicated that the members shown in Table 3 are sufficient under present train loadings and track speeds. The information developed from this study on the structural behavior of the Stoney Creek bridge under modern train loadings will enable the strategic planning of a rehabilitation program that will ensure the safety of train operations for future railway traffic over Stoney Creek Bridge. ACKNOWLEDGEMENT The writer wishes to acknowledge Mr. D.P. Gagnon, P.Eng. of Buckland and Taylor Ltd. and Mr. D.E.J. Adamson, P.Eng. of Canadian Pacific Railway for their assistance in the preparation of this paper.

12 REFERENCES American Railway Engineering and Maintenance-of-way Association (AREMA 1998). Manual of Recommended Practice: Steel Structures: Existing Bridges, Chapter 15, Part 7. Washington, D.C. Ammann O.H. (1918). The Hell Gate Arch Bridge and Approaches of the New York Connecting Railroad over the East River in New York City. Transactions of the American Society of Civil Engineers 82, New York. Buckland and Taylor Ltd. (1999 & 2000). Stoney Creek Arch Bridge Load Rating. North Vancouver, B.C. Canadian Pacific Railway (CPR 1992). Coal Route Fatigue Study. Montréal, QC. Conway W.B. (2001). Practical Application of the Rating Rules AREMA Annual Conference, Chicago, IL. Cook R.D. (1981). Concepts and Applications of Finite Element Analysis. J. Wiley, New York Dick, S.M. & McCabe, S.L. (1991). Improved Techniques for Evaluation of Railway Bridge Fatigue, 8 th Annual International Bridge Conference, Pittsburgh, PA. Engineering News (1894). Stony Creek Arch, Canadian Pacific Railway. Engineering News August 2. New York. Lavallée O. (1974). Van Horne s Road. Railfare Enterprises, Montréal, QC. Motley P.B. (1930). Reinforcement in Place of the Stoney Creek Arch Bridge. Engineering Journal Vol. XIII, No. 5. Montréal, QC.

13 LIST OF TABLES TABLE 1. Number of Deficient Bridge Members at Normal Rating TABLE 2. Governing Deficient Members of Each Member Type TABLE 3. Maximum Cooper s Axle Load for D/C = 1.0 LIST OF FIGURES FIGURE 1. Stoney Creek Bridge FIGURE 2. General Elevation, Plan and Design Wheel Load for 1893 Steel Arch FIGURE 3. General Elevation, Cross Section and Design Wheel Load for 1929 Steel Arch FIGURE 4. Three Dimensional Model of Arch FIGURE 5. Deficient Members of Truss Arch

14 TABLE 1. Number of Deficient Bridge Members at "Normal" Rating 1893 Truss Arches 1929 Truss Arches TC (32) BC (32) V (34) TC (32) BC (32) V (34) number in parenthesis denotes total number of members of that type in the bridge TC=Top Chord; BC=Bot. Chord; V=Verticals

15 TABLE 2. Governing Deficient Members of Each Member Type Combined Forces from 3D Analysis (Demand) Member Axial (kip) M LONG (kip-ft) M LAT (kip-ft) D/C at 60 mph Max. Speed for E-80 & D/C=1.0 (mph) Truss Arch Top Chord Members TC (4)* TC (4)* TC (4)* not possible TC (2)* TC (4)* TC (2) TC (2) not possible Truss Arch Bottom Chord Members BC (3)* BC (4)* not possible BC (1)* BC (4)* not possible BC (4)* not possible BC (4) not possible BC (2) not possible BC (4) not possible BC (2) Truss Arch Verticals P (1)* P (1) P (2) not possible Spandrel Bent Members Transverse Beam P Deck Plate Girders 26 ft span at west ft span at west ft span (P0 - P1)* ft span (P1 - P2)* number in parenthesis denotes total number of deficient members of that type in the bridge, values shown in the table are those of governing member * members with governing D/C at west end of the bridge subject to track curvature

16 Table 3. Maximum Cooper s Axle Load for D/C = 1.0 Member D/C Cooper s Axle Load (@ 60 mph) TC (4) 1.21 E-64 TC (4) 1.32 E-59 TC (4) 1.10 E-73 TC (2) 1.14 E-70 BC (4) 1.22 E-60 BC (4) 1.18 E-64 BC (4) 1.23 E-61 BC (4) 1.14 E-67 BC (2) 1.14 E-67 BC (4) 1.14 E-67 P (2) 1.21 E-64

17 FIGURE 1. Stoney Creek Bridge

18 FIGURE 2. General Elevation, Plan and Design Wheel Load for 1893 Steel Arch

19 FIGURE 3. General Elevation, Cross Section and Design Wheel Load for 1929 Steel Arch

20 FIGURE 4. Three Dimensional Model of Arch

21 FIGURE 5. Deficient Members of Truss Arch 25 mph)