Test Results for the Fatigue Design/Rating of Thru Truss Hangers Robert A. P. Sweeney, PhD, D. Eng., P. Eng., Eng.

Former Chief of Structures, Canadian National/Illinois Central

Modjeski & Masters

4675 Doherty

Montreal, Qc H4B 2B2

Canada

Tel/Fax: 514-483-4935

ABSTRACT This paper deals with the design and evaluation of hangers on through trusses for fatigue, and discusses the results of recent field test of hangers at CN.

The paper also includes a review of recent corrections to the methodology of handling

tension members in Chapter 15 (1).

It is demonstrated that the current AREMA provisions for calculated allowable stress

ranges for hangers are reasonable from a statistical perspective (95th percentile).

Impact for design covers a rare occurrence that might occur only several times during the

design life of a structure (4). Test results will be reported that show that the design

formula for impact may be reduced to 40% of the calculate value used to design hangers

for maximum load, again within reasonable statistically certainty (95th percentile).

It will be shown that further reductions in stress may be found through testing of

individual hangers.

An example of a hanger fatigue failure is discussed, together with an example of a

potential pin plate failure.

INTRODUCTION

Hangers have been a problem for many years, and in most through trusses their

connections are the most failure prone members of the main truss system, excluding

members failed through accidental impacts caused by collisions. It is difficult to predict

the stress distribution at hanger connections; hence it is even more difficult to predict

their useful fatigue life.

AREMA Chapter 15 (1) is going through a major change in concept with regard to

tension members that requires a thorough re-think of long standing assumptions.

Early analysis methods, using plane truss analysis, ignored the very real stresses induced

by in and out-of-plane bending. Friction in pinned joints and plate rigidity in riveted

joints caused this together with in and out-of-plane forces and/or loads.

There are many papers covering and evaluating this with very important work done at the

University of Illinois in the 1940’s and 1950’s and summarized by AREA at the time (2).

In the early days of hanger design, with design loads that exceeded actual loads by a large

margin, the differences for design purposes were covered by ignoring the secondary

stresses (See current Article 15.1.3.15).

At some time, it became evident as a result of failures that this was not sufficient and the

allowable design stresses were reduced to 70.7% of the normal allowable tension. This is

equivalent to allowing for a 41.4% increase in stress due to bending effects.

Michalos’s paper (3) published by AREA in 1957 gave a hand calculation method for

obtaining the out-of-plane stresses, and any number of plane frame analysis methods

would get the in-plane bending stresses assuming full joint fixity.

If these methods were strictly followed it should work for the failure case of Ultimate

failure with the appropriate safety factors of either Normal or Maximum Rating.

For failure due to fatigue that occurs at much lower stresses and is controlled by stress

range as opposed to maximum stress, the problem is to predict when failure is a potential

concern. Few railroads have the luxury of being able to restrict the weight of rolling

stock to ensure that fatigue of these members will never be an issue (so-called Fatigue

Rating), and the Manual does not recommend such a limitation.

The objective is to predict the when, given the load, and not how much load. Given the

normal or maximum rating (depending on the line usage and importance and future

capital spending plans), the objective is to determine when there is significant risk of

failure caused by fatigue.

Over time the methodology has evolved and prediction are getting much better.

But first a review of general tension member behavior at the ultimate is necessary.

GENERAL TENSION MEMBER BEHAVIOR

Since its inception, the AREA Recommended Practice and subsequently the AREMA

Recommended Practice has based the limit on tension members as a function of the yield

stress and the net area. Concern for yielding and distortion were paramount. At one time

there was an additional limit that the allowable used could not be greater than half the

ultimate tensile limit (0.5 Fu). Over the years this was dropped.

When the margin between the yield stress and the ultimate tensile strength was large this

was an acceptable procedure. With today’s modern high strength steels that are now

accepted in Chapter 15, the margin between the yield and ultimate tensile strengths is not

large, and additionally, the current AREMA Guidelines (2003) in Chapter 15 are

misleading as to the appropriate failure modes.

The proper criteria is to apply the gross area to the limiting yield stress and the net area to

the ultimate tensile stress both with an appropriate margin of safety or error, and of

course use the lower of these.

Letter ballots are being evaluated to hopefully include in the 2004 edition of the Manual

the following:

Proposed Manual Revisions

For members subject to tension, the stress shall not exceed:

0.55Fy on the Gross area, and not the net area, and

0.5 Fu on the net area except for pin connected members where the allowable is 0.45Fy at

the location of the pinhole.

Net area is calculated with the s2/4g reduction (Article 15.1.5.8) and may be further

reduced to account for shear lag (Article 15.1.6.5).

Tension in floor beam hangers, including bending based on the net section is limited to

14,000 psi (0.707 x 19,800 psi) for riveted end connections and 19,800 psi (0.55 x 36,000

psi) for high strength bolted connections. This is also under review.

In beams, the limit for tension in the extreme fibers based on the net section is 0.55 Fy.

Fatigue criteria

For fatigue, the appropriate section to use depends on the connector. For high strength

bolted connections and welded connections it is the gross section, and for riveted

connections, eyebars and pin plates it is the net section.

These differences are based on the history of how the test results were reported and

sorting out which is correct remains a task for the future. At low stress ranges typical of

railroad bridge loadings and at small crack sizes, the behavior is probably best described

with reference to the gross section. But, at large crack sizes where the piece is about to

fail in a brittle manner, the net section probably applies.

Typical Hanger Connections

If one considers a hanger which typically is connected by two plates, one on the outside

and one on the inside (track-side) of the hanger, how is the fact that the full section is not

engaged to be applied to fatigue?

Not a problem with net section calculations, but what about gross section calculations?

Where will the failure occur?

The complication is the fact of shear lag in tensile connections where the entire

connection is not directly connected. This concept is really only meant to apply to the net

section calculation at ultimate (working stress at allowable limit) and has no real meaning

with regard to failures at lower stress ranges under a fatigue regime. It has become

practice to use the shear lag factors for fatigue calculations where the appropriate section

is the net section.

When the appropriate section is the gross section, the correct answer is somewhere

between the gross section reduced by a calculation similar to the shear lag reduction and

simply using the parts actually connected. The first has no basis in test results and the

second is conservative.

Due to a complete lack of test results, AREMA Chapter 15 will be recommending the

later for gross section calculations on high-strength bolted hanger connections, although

this is unlikely to appear before the 2005 Edition.

Actual fatigue evaluation calculations follow the procedures outlined in Chapter 15, part

7.3.4.2. Note that special stress ranges for Fracture Critical members will be eliminated

in the 2004 edition of the AREMA Manual.

CN TESTING

At CN, a testing program aimed at determining the parameters to aid in this remaining

life calculation of many members were investigated by full scale field testing after a rash

of near failures that were prevented by the Company’s inspection program and perhaps a

lot of luck.

With regard to hangers, the conclusions of these studies were:

• In and out of plane bending effects on hangers were safely covered, except for

one point, by the 1.42 times the axial stress criterion or reducing the allowable to

0.707. Clearly an analysis of the actual bending effects, either analytically or by

field measurement, would allow for a more precise estimate of stress ranges. The

current reduction giving actual stress limits is equivalent to a 0.707 reduction

(1/1.42) and is based on A36 material, but such low yield material is rarely used

in current designs. This criteria needs to be reviewed. See Figure 1.

• Although in many instances, the measured stress ranges and stresses in field tests

are less than those predicted, it was found that to be safely covered (95th

percentile) it was not prudent to reduce the calculate stresses unless proper field

testing confirmed such a reduction actually occurred. Clearly, the probability that

lower stresses would be measured than those calculated is high but never

guaranteed. Hence, the Manual recommends that the alpha factor

(measured/calculated maximum stress) be taken as 1.0 unless field measurements

confirm otherwise.

• Impacts based on the design formulae in the AREMA manual are expected to

occur 3 to 4 times at most in the life of most structures. A study of the impact test

results indicated a probability of occurrence of no more than 1% (4).

• Studies at CN, which has the largest data base of test results for typical fatigue

stress ranges ever reported, indicate that using 40% of the AREMA Calculated

Impact formula would safely cover (95th percentile) the effect of this parameter on

the actual stress ranges in hangers on Class 1 good track with good wheel control.

• On a floor beam spacing of 22 feet, and Truss spacing of 18 feet, this would result

in a mean impact for fatigue calculations of 0.4 x 44.6 % = 17.6%, where 44.6%

is the AREMA calculated impact for that loaded length and truss spacing. Again,

actual field measurement of the impact has a good chance of showing a lower

value, but cannot be relied upon for an individual structure without a field test.

See Figure 2.

Using the above information to predict useful remaining fatigue lives of hangers gives a

better estimate of the first indication of when there should be concern. Depending on the

economic and capital plan of the railroad concerned, it is possible to refine these

calculations considerably by field-testing to get a more precise and generally a longer

remaining calculated fatigue life since both these parameters a likely to be less severe

than the assumptions recommended above.

Test for actual alpha factor (Figure 1).

This requires a static test with the applied load located at the maximum and minimum

points on the influence line. A slow crawl test with a heavy vehicle(s) loading the full

influence line and ensuring minimum dynamics effects is often used due to time

constraints. Figure 1 plots the actual vs. calculated maximum stresses based on net

section for riveted and pinned connections.

Alpha Hanger Data

0

2000

4000

6000

8000

10000

12000

14000

16000

18000

20000

0 2000 4000 6000 8000 10000 12000

Simple Truss Theory psi

Test

Dat

a

Data Simple TheoryA + 1.65 Sigma (1.62)1.33 * Theory (1.22)1.42 * theory (1.14)

Figure 1

Test for actual stress ranges.

This would require typical trains run at the most probably speed used at the location and

are not reported here, although such tests were carried out.

Test for typical average impacts (Figure 2).

Impact Truss Hangers over 10 mph

0

5

10

15

20

25

30

35

40

45

50

0 5 10 15 20 25 30 35 40 45 50

Theory Impact

Mea

sure

d Im

pact

Test DataTheoryAvg + 1.65 Sigma (0.374)

Figure 2

This can be done with a test rain run at various speeds to get impact data. Figure 2 shows

the critical data from tests on 13 hangers. The 95th percentile is at 37.4%. The

committee rounded this to 40%. All of theses tests were conducted with good track

conditions and good wheel conditions typical of top quality Class 1 railroad track and

wheel control.

Other More Refined Techniques

With calculations based on strain measurements and actual as opposed to assumed impact

percentages, it is possible to further refine the calculations using fracture mechanics

analysis.

When this is all done, it is still possible, if it make economic sense, to make use of

acoustic emission testing to verify crack movement and of course good inspection is

critical in dealing with these members.

It is this writer’s belief that more than visual inspection is required and recommends

magnetic particle inspection of suspect hanger connections.

FATIGUE : EXAMPLE OF HANGER FAILURE

The truss shown in figure 3 is typical of many riveted trusses with hangers connected to

the top chord and end post by plates on the track-side and water-side only.

Several years prior to the failure, based on fatigue life calculation done internally at CN,

this truss and several others were put on a watch list requiring detailed inspection, but the

watch was put on the bottom connection to the floor beam and not on the top connection

to the top chord and end post.

Figure 3 Overall view hangers

Strains were measured in the temporary replacement hangers after failure, and using

projection, the strains were estimated at the rivets where the section failed and a rough

after the fact fatigue analysis was made indicating failure had been imminent. The strains

were converted to stresses at the failure location using the net section area reduced by the

standard shear lag calculation, and compared to the limits for reamed rivet holes shown in

Figure 15-9-4 of Chapter 15 of the AREMA Manual that is the appropriate fatigue curve

for the resulting root-mean-cube stress range.

Figure 4 Failure Surface in Field

Figure 4 shows the rupture of the piece in the structure after it was re-connected with a

temporary bolted connection. Slow fatigue crack growth in the outstanding leg of the

angle starting in the rivet hole in that leg that turns and grows along the inside leg of the

angle to the rivet hole connecting the batten plate. At this point rapid failure occurred.

There were also fatigue cracks in the other outstanding legs as seen in Figure 5.

Figure 5 Matching Failure Surface in Lab

Figure 5 show the matching piece in the lab. The top photo shows the piece upside down

from what it would have been in the field. The bottom photo shows the piece with the

outstanding angle legs on the table.

The failure stared with a fatigue crack that initiated at the outside hole in the outstanding

leg of the waterside angle. This crack progressed slowly until it popped through the rivet

hole connecting the angle to the batten plate at which time the rapid rupture occurred.

Note the slow crack growth on the piece on the left in the bottom photo as opposed to the

tearing on the right hand piece.

Figure 6 Fatigue to brittle fracture and tearing

On the left side in the top photo of figure 6 the failure surface is a lot smoother,

indicating fatigue crack growth, than the right side where tearing is evident.

As a result of this failure, which caused considerable damage to the floor system stringer

connections and the opposite floor beam to hanger connection, a special inspection was

arranged for all similar trusses on the line. All had been designed by the same firm and

installed within a few years of each other.

Cracks were found in most of these trusses similar to the one shown in figure 7, bottom

photo. The top photo shows the connection. The bottom photo shows a close up of the

Figure 7 Cracking before failure Typical

waterside gusset connection. Note how the crack that is just above the yellow line is just

entering under the rivet connecting the outstanding angles to the batten plate. When the

crack pops out the other side of the rivet hole, the piece will fail in a spectacular manner.

There were instances on other trusses where the cracks initiated on the inside as opposed

to the outside.

A retrofit program was arranged which consisted of adding an additional plate bolted on

the inside and outside of the top connection similar to what is shown in figure 8.

Figure 8 Temporary Repair

POTENTIAL PIN-PLATE FAILURE

Office calculation indicated a potential fatigue problem with the pin-connected hangers

on the following structure (6,7).

In 1975 (6,7) and again in 1979 these hangers were strain gauged. Special stress

concentration gages (measuring 10 strains over a very short distance) were placed as

close to the pins as possible at the potential critical location on the pin plates to ascertain

the stress concentration factors. The results correlated well with the “Kobyashi” equation

used to estimate the maximum stress concentration (5, 6).

Figure 9 Inside Pin Plates

Calculations based on these strain gage measurements predicted potential failure by 1984.

The joints were retrofitted by 1982.

Figure 10 Pin Plate after 75 years

Figure 10 and 11 show a hanger pin plate from a companion span that was knocked down

and destroyed by a barge in 1976 (6,7).

Figure 11 Close up Forging marks or cracks? If the pin plate in figure 11is viewed as a clock, cracking would be expected first at

between 2 and 3 o’clock positions or the 9 to 10 o’clock positions. There was sufficient

concern as to whether these were forging marks or cracks, and if forging marks could

they lead to serious cracks.

The precise location of potential cracking in the pin plates is highly influenced by friction

around the pin.

Given that the structure was intended to remain in service for many years to come, it was

decided to re-place the pin plates at the four remaining critical sub-hangers with a high-

strength bolted retrofitted connection. Figure 12 shows the retrofitted hanger

connection.

Figure 12 Bolted Retrofit Connection

CONCLUSIONS The paper outlines that for fatigue evaluation:

1 Calculated static stresses must be assumed to be as calculated unless field full

scale load tests on the member being evaluated show lower values,

2 Impacts may be assumed to be 40% of the AREMA Calculated values for Class 1

mainline track and wheel control unless similar tests on the actual member being

evaluated show lower values, and

3 For both parameters, the possibilities of measuring values that are lower are high

making testing a worthwhile exercise.

The change in philosophy on how to deal with tension members is outlined emphasizing

that gross section is appropriate for yield calculations and net section is appropriate for

ultimate tensile strength calculations while pointing out some inconsistencies due to the

way past test results have been reported.

Finally an example of a hanger riveted connection fatigue failure is illustrated, and a

potential pin connection failure is also shown.

ACKNOWLEDGEMENT

The writer acknowledges the staff at CN who assisted him in his 37 years with the CN-IC

System together with Modjeski & Masters Inc. for their continued support and

encouragement.

REFERENCES

1 American Railway Engineering and Maintenance of Way Association Manual of

Recommended Practice, Chapter 15, 2003, AREMA, Landover, MA

2 Committee on Iron and Steel Structures, Stress Distribution in Bridge Frames –

Floorbeam Hangers, Proceedings AREA Vol. 51, 1950, pp. 470-503

3 James Michalos and J.M. Louw, Vol. 58, pages 1 - 51, AREA Proceedings 1957

4 Byers, William G., “Impact from Railway Loading on Steel Girder Spans”, ASCE

J. of Structural Division, Vol. 96, No, ST^, June 1970, pp 1093 – 1103

5 Fisher, J.W., “Fatigue and Fracture in Steel Bridges – Case Studies” J. Wiley

1984, particularly page 32.

6 Sweeney, R.A.P., “Load Spectrum for Fraser River Bridge at New

Westminster, B.C.”, Proceedings A.R.E.A., Volume 77, Bulletin 658,

June-July 1976, A.R.E.A., Washington, D.C.

7 Fisher, J.W., Daniels, J.H., “An Investigation of the Estimated Damage in

Members of the 380 ft. Main Span, Fraser River Bridge”, Proceedings A.R.E.A.,

Volume 77, Bulletin 658, June-July 1976, A.R.E.A., Washington, D.C.