TECHBRIEF

Research, Development, and 

Technology

Turner-Fairbank Highway 

Research Center

6300 Georgetown Pike

McLean, VA 22101-2296

www.fhwa.dot.gov/research

Guidelines for Design and 

Rating of Gusset-Plate

Connections for Steel 

Truss Bridges

FHWA Publication No.: FHWA-HRT-14-063

FHWA Contact: Justin Ocel, HRDI-40, (202) 493-3080, 

justin.ocel@dot.gov

This document is a technical summary of the published 

National Cooperative Highway Research Program 

Web-Only Document 197, Guidelines for the Load 

and Resistance Factor Design and Rating of Riveted, 

Bolted, and Welded Gusset-Plate Connections for Steel 

Bridges, available through the Transportation Research 

Board at http://www.trb.org/Main/Blurbs/168923.aspx.

INTRODUCTION

On August 1, 2007, the I-35W bridge collapsed in 

Minneapolis, MN. The bridge was a three-span (255 by 456 

by 255 ft), continuous subdivided Warren deck truss over the 

Mississippi River. The National Transportation Safety Board 

(NTSB) investigated the collapse, with technical assistance 

from the Federal Highway Administration (FHWA). NTSB 

determined that the collapse was due to undersized gusset 

plates at a particular node location on the main truss. The 

plates should have been twice as thick or should have had 

twice the yield strength to provide adequate resistance to 

prevent the collapse.(1)

During the failure investigation, it was clear that the failure 

of I-35W was purely attributable to the gusset plate design. 

Recognizing that there could be lingering safety issues with 

other gusset plates in the bridge inventory, NTSB issued the 

following recommendation to FHWA on January 15, 2008:

“For all non-load-path-redundant steel truss bridges 

within the National Bridge Inventory, require that 

bridge owners conduct load capacity calculations to 

verify that the stress levels in all structural elements, 

including gusset plates, remain within applicable 

requirements whenever planned modifications or operational changes may significantly increase stresses.”(1)
2

On the same day, FHWA issued Technical 

Advisory (TA) 5140.29, which strongly encouraged bridge owners to check the capacity of 

gusset plates in their inventory.(2) At the time, 

bridge load ratings were member-centric and 

did not require connection elements to be 

rated, but the TA describes a paradigm shift 

in load rating. Based on feedback from many 

sources, it became apparent that there was 

no consensus on the specific procedures to 

follow for design or rating of gusset plates. 

At the time, the American Association of State 

Highway Transportation Officials (AASHTO) 

Bridge Design Specifications (BDS) were vague 

on the subject, leaving room for considerable 

engineering discretion and judgment in the 

process.(3) To provide a uniform standard for 

load rating gusset plates, FHWA issued a guidance document in February 2009, based on 

the best available information on gusset-plate 

behavior, outlining the minimum number of 

resistance equations that must be checked to 

adequately load rate a gusset plate.(4) For this 

TechBrief, this document will be referred to as 

the FHWA guide.

DESCRIPTION OF THE FHWA GUIDE

The FHWA guide provides rating guidance 

on both load and resistance factor rating 

(LRFR) and load factor rating philosophies. 

Discussions in this TechBrief are from the LRFR 

perspective only. The FHWA guide recommends five resistance checks at a minimum. 

First, fasteners must be checked for adequate 

shear resistance for each connected member. In addition, the bearing resistance of 

the plate material must also be checked as 

part of the fastener checks. The FHWA guide 

refers to existing resistance checks in the load 

and resistance factor design (LRFD) BDS.(3)

To assist with rating older riveted bridges, 

the FHWA guide also provides recommended 

shear strength values for rivets based on their 

age because this information is not provided 

in recent AASHTO specifications. 

Second, relevant planes through the gusset 

must be checked for shear yielding and shear 

fracture. The one caveat associated with the 

shear yielding limit state is a suggested multiplication of the resistance by a factor, Ω, 

which the engineer has discretion to choose 

as 1.00 or 0.74. The  Ω-factor was first introduced in 1985 by the authors of the AASHTO 

Guide Specifications for Strength Design of 

Truss Bridges and accounts for assumptions 

in the shear stress distribution.(5) The specification writers recognized two potential outcomes: (1) the gusset plate would behave 

as a beam in bending along a shear plane, 

resulting in a parabolic distribution of stress 

(Ω=0.74), or (2) if adequately stiff or braced, 

the gusset plate would redistribute yielding 

along a shear plane until the entire section 

was mobilized (Ω=1.00). The original authors 

believed that gusset plate behavior could fall 

anywhere within these two bounds. Therefore, 

the FHWA guide left the selection of the factor 

to the discretion of the load-rating engineer. 

Finally, the ɸ-factors for tension limit-states 

were applied to the shear limit-state checks, 

these being 0.95 for shear yielding and 0.80 for 

shear fracture. These values were set knowing 

that, for ductile materials, shear failures are 

technically tensile failures on inclined planes.

Third, the gusset plate must be checked for 

local stability around compression members. 

The buckling limit of the gusset plate is calculated by treating the zone of plate between the 

end of the compression member and adjoining 

members as an equivalent column and using 

the column formulas in the LRFD BDS to calculate capacity. This is done by defining an effective area over which the member axial load is 

assumed to be distributed. This area is referred 

to as the Whitmore section because it is based 
3

on the writings of Whitmore.(6) According to 

the Whitmore theory, the stress spreads out 

from the member into the gusset plate by 

dispersing at 30-degree angles from the first 

to the last row of fasteners. The width of the 

section is defined along the end row of fasteners and extending out to the dispersion lines. 

Finally, an equivalent column length is defined 

as the average of three lengths projected from 

the Whitmore section, in the direction of the 

member, to the fastener lines of the adjoining members. The three lengths are taken at 

the two ends of the width and at the center. 

If the Whitmore section intersects an adjoining member bolt line, the length at that end 

is assumed to be zero. This is schematically 

shown in figure 1. The load-rating engineer is 

given leeway to select an appropriate equivalent column length, or K-factor, depending on 

the likely buckling mode shape. Suggested 

ranges of K are given based on the rotational 

restraint at the ends and whether the gusset 

plate has lateral support to prevent sway. 

Fourth, for tension members, the same 

Whitmore section is defined at the end of 

the tension members. On this section, both 

gross yielding and net section fracture must 

be checked. In addition, the fifth check is block 

shear resistance around tension members.

STATEMENT OF WORK

Early in the implementation of the FHWA 

guide, some gusset plates were not attaining desirable rating factors during load rating, leading States to retrofit or restrict loads 

on those bridges. Because many of these 

bridges had no long-standing history of performance problems, the common thought 

was the FHWA guide was overly conservative, 

particularly for fastener shear, gusset shear 

yielding, and Whitmore compression checks. 

AASHTO (through the National Academies), 

along with FHWA, sponsored an experimental 

program to further enhance the understanding 

of gusset plate failure mechanisms and create 

refined resistance equations that could predict 

the various failure modes of gusset plates. 

This project was the National Cooperative 

Highway Research Program (NCHRP) Project 

12-84, Guidelines for the Load and Resistance

Figure 1. Example of Whitmore section derived from 30-degree dispersion angles.
4

Factor Design and Rating of Riveted, Bolted, 

and Welded Gusset-Plate Connections for 

Steel Bridges.

(7)

The research included both experimental testing and analytical modeling. Primarily, the 

physical limitations and expense of experimental testing dictated that a small sampling 

of specimens would be tested to provide an 

adequate number of finite element model 

calibration points. Once robust modeling 

procedures were established, using the 

experimental results as a benchmark, a much 

broader study of different connection geometries could be investigated analytically to 

encompass the types of gusset plates that are 

in the Nation’s inventory of truss bridges. The 

project goal was to collect the data to support, 

refute, or expand on the load rating checks 

in the FHWA guide. The data could then be 

adopted into the AASHTO BDS and Manual for 

Bridge Evaluation (MBE).(8,9)

EXPERIMENTAL PROGRAM

The experimental program specifically tested 

12 full-scale gusset plate connections. The 

members were reusable, and each new specimen consisted of two new gusset plates and 

a set of chord splice plates. The configuration of the experimental connections used five 

separate members: two collinear chords, one 

compression diagonal framing at 45 degrees 

to the chord, one tension diagonal framing 

at 45 degrees to the chord, and a vertical 

member that could be in either tension or compression framing perpendicular to the chords. 

Figure 2 illustrates the special-purposed load 

frame constructed to test the gusset plate 

connections. Four of the five members had 

independent control of load via servovalvecontrolled hydraulics. This setup offered the 

ability to investigate different equilibrium load 

combinations on the individual specimens.

Figure 2. Load frame used for experimental testing. (Units shown in inches.)
5

The testing was conducted in two phases. The 

first phase established six unique geometries 

of plates. Different bolt strengths were investigated, along with the free edge length around 

the compression diagonal and span length of 

plate at the end of the compression diagonal. 

The plates tested had dimensions of roughly 

60 inches wide by 36 inches tall. The phase 

2 specimens used the same geometries but 

had modified plates with artificial section loss, 

edge stiffening, and multi-layered gussets.

PARAMETRIC PROGRAM

Analytical models of the specimens helped 

define the level of detail needed to predict the 

experimental failure with accuracy. A threedimensional shell model of the gusset plates 

and members was necessary to properly 

predict the failure of the connection. For the 

purposes of determining shear yielding and 

buckling, fastener holes did not need to be 

modeled. However, nonlinear material and 

geometric properties of the gusset plate were 

necessary, along with initial geometric imperfections.

The robust finite element modeling technique was applied to a parametric study that 

increased the breadth and depth of studied 

connection geometries over the experimental 

study. In particular, this included connections 

with diagonals framing at angles other than 

45 degrees; chords that were not collinear; 

corner joints; Warren and Pratt configurations; 

loading scenarios representing joints over a 

pier, at midspan, and near inflection points 

of trusses; gusset plates with edge stiffening; 

gusset plates with section loss; and multilayered gusset plates. In total, there were 201 

different models. Owing to the fidelity of the 

models, only buckling, shear yielding, and 

chord splice failure modes could be identified. 

The models did not have the fidelity to capture 

net section type failures.

SELECTED RESULTS

Because many bridge owners had begun 

to load-rate gussets according to the FHWA 

guide while the research was ongoing, the 

conclusions and recommendations focus on 

improving and expanding the methodology as 

already defined. The following sections highlight the important results identified in the 

research.

Shear

In total, there were 44 observed shear-yielding 

failures from the experiments and analytical 

models. On average, the real shear-yielding 

load was 2 percent higher than the predicted 

0.58FyAs value used in the FHWA guide. This 

would imply, in a nominal (unfactored) sense, 

that Ω should be 1.00. However, to account for 

the scatter in the data, Ω was selected as 0.88 

for reasons to be explained later.

Buckling

The analytical modeling matrix included the 

geometry at the ends of compression diagonals as a variable. In reviewing many truss 

plans, members often had chamfered ends 

so they could be spaced as close as possible 

to adjoining members. However, there was 

a substantial difference in buckling strength 

between chamfered compression members 

and those that were not. This contrast is highlighted in the stress contours in figure 3 and 

figure 4, which show similar connections with 

a non-chamfered and a chamfered diagonal, 

respectively. When the compression diagonal was not chamfered, buckling occurred 

with very little yielding, whereas when the 

chamfer was present, buckling still occurred, 

though after a significant amount of yielding. 

Closer analysis found that shear dominated 

the yielding along the chamfered edge. Hence, 

for joints with closely spaced members, the 

buckling strength was found to be coupled to 
6

the shear strength of the planes around the 

compression member. However, these shear 

planes are not across the full width or height 

of the plate and thus are referred to as partial 

shear planes. This led to a two-folded resistance check to determine buckling resistance. 

First, the Whitmore buckling resistance had to 

be evaluated, and the shear yielding on a partial plane around just the compression member also had to be checked. Figure 5 illustrates 

potential partial planes.

The FHWA guide check requires an effective column length based on the average 

of three lengths along the Whitmore width 

as described earlier. The work conducted 

throughout the research period found a better 

correlation when the effective column length 

was based only on the length at the middle of 

the Whitmore width, rather than the average 

of three lengths. In addition, a fixed K-factor 

of 0.5 was found applicable for all gusset 

plates in predicting the minimum buckling 

resistance.

In total, 124 experiments and analytical models 

were evaluated with the combined Whitmore 

buckling and partial plane shear yield checks 

to define the statistics to be used in the resistance factor calibration. In general, thin plates 

(3

/8 inch thick or less) were controlled by the 

Whitmore buckling criteria, and thick plates 

were usually controlled by the partial plane 

shear yielding criteria. 

Chord Splice

The FHWA guide suggests that Whitmore 

buckling and tension checks should be used 

to evaluate compression and tension chord 

splices.(4) In the analytical models that failed 

in the chord splice, the stress patterns did 

not match those of the assumed Whitmore 

Figure 3. Von Mises stress contours with no chamfering of compression diagonal end. (Grey contour represents yielded 

material.)
7

section, in addition to the ambiguity of load 

sharing between the gusset plate and any 

splice plates. The research identified a resistance check that treated the gusset and splice 

plates as a single cross-section and considered the axial and bending component from 

the eccentric chord loads. 

In total, 33 analytical models were observed 

to fail in the tension or compression chord 

splice. There were no experimental failures of 

chord splices. 

Block Shear

This research specifically did not address 

the block shear limit state because there are 

abundant experimental data available in the 

open literature. The best summary of the available block shear data can be found in Huns, 

Grondin, and Driver.(10) This paper summarizes 

133 experiments reported from 8 different 

sources. The paper provides the statistics 

needed for resistance factor calibration using 

the current AASHTO block shear equation.(8)

Other Tension Failure Modes

Only three analytical models had tension diagonal and/or vertical members that demonstrated pure tensile failure modes. All other 

models or specimens had tension members in 

conjunction with compression members that 

failed in buckling or shear. Therefore, there 

were no data available to provide a statistically significant sample to derive changes 

for tensile failure modes, so no changes were 

recommended to the FHWA guide for 

Whitmore tension checks. 

Edge Stiffening

No correlation could be identified between 

the buckling resistance of a gusset plate and 

Figure 4. Von Mises stress contours of model with chamfered end compression diagonal. (Grey contour represents 

yielded material.)
8

the free-edge slenderness of the plate itself. 

If properly detailed, edge stiffening could 

be used to increase the buckling resistance 

of a gusset plate, and this was proven both 

experimentally and analytically. In this case, 

properly detailed stiffeners would add out-ofplane stiffness to the compression member. 

Conventional edge stiffeners stiffen only the 

free edge and therefore offer no additional 

gains in buckling resistance. If stiffened properly, the Whitmore buckling criteria could be 

ignored, but the partial plane shear yielding 

resistance would still cap the overall resistance of the compression member.

No changes to the existing edge slenderness 

limit in the BDS were recommended because 

it represents good detailing practice to assist 

with transportation and erection. However, 

additional commentary was recommended to 

explain that it is not meant to suppress buckling. It was recommended not to rate gusset 

plates based on edge slenderness.

Corrosion

The effects of corrosion were investigated both 

experimentally and analytically, by applying 

localized areas of section loss. For shear limit 

states, it was found that resistance was based 

on the total cross sectional area in the plane of 

failure. Therefore, there was no effect from section loss being unbalanced across the gusset 

plate or unbalanced between the two plates. 

Evaluation of buckling with section loss followed a similar reasoning. In this case, any 

section loss occurring between the Whitmore 

width and the adjoining member fastener lines 

is projected onto the Whitmore section, and 

an equivalent plate thickness is derived on 

the Whitmore section. This equivalent plate 

thickness is used in the resistance equations. 

Overall, this approach was found to be very 

conservative at times, but there were not 

enough failures to derive any better recommendation.

Figure 5. Partial Shear planes around compression member.
9

Multilayered Gusset Plates

For observed failures involving multilayered 

gusset plates, it was appropriate to add the 

individual resistances of each plate together 

to attain the overall resistance. However, consideration should be given to determining 

whether enough fasteners are being provided 

to develop the strength of each plate at the 

plane of failure.

DESIGN AND EVALUATION 

SPECIFICATIONS 

Resistance factors were determined through 

Monte Carlo reliability analysis according to the 

philosophy outlined in NCHRP 20-07(186).(11)

To expedite the calibration process without 

conducting a series of live-load analyses 

through a suite a various trusses, it was easier to assume a dead-to-live load (DL/LL) ratio. 

Twenty different truss construction plans, from 

eight different States, were used to define the 

variation in DL/LL. The DL/LL ranged from 0.5 

to 6.5, with an average of approximately 2.0.

The AASHTO Strength I and IV load combinations were used in the Monte Carlo analysis. 

Generically shown in figure 6 is the variation 

of the resistance factor, ɸ, with the DL/LL ratio 

using the two AASHTO load combinations. 

The figure shows required resistance factor 

decreasing with the DL/LL for the Strength I 

load combination and increasing with DL/LL 

with Strength IV. Typically, the two would converge at a DL/LL between 6.0 and 8.0. The solid 

black line in the figure shows what the overall 

resistance factor should be between the two 

distributions. However, final resistance factors are defined at discrete DL/LL ratios for the 

Strength I load combination, rather than having a curve of many different ratios.

Figure 6. Sample ɸ-factor variation with DL/LL ratio.
10

The results from this research made a distinction between design and rating, because the 

greater challenge is rating the existing inventory. It is more expensive to retrofit an existing 

gusset plate with an inadequate rating than 

to increase the plate thickness in design. It 

was out of scope for this project to define a 

target reliability for gusset plates. Therefore, 

a decision was made to base design resistance factors on the higher reliability index of 

4.5, whereas AASHTO BDS uses a target reliability of 3.5. There were two reasons for this 

decision: 1) AASHTO already requires a higher 

level of reliability for fastener, weld, and block 

shear limit states, and 2) the American Institute 

of Steel Construction Steel Construction 

Manual uses two reliability indices (2.8 for 

member-centric limit states and 4.0 for connection-centric limit states).(12) Because gusset 

plates are a connection element, it seemed 

prudent to calibrate to a higher level of reliability. 

The scatter in the resistance predictions was 

sensitive to the gusset plate thickness, so different calibrations were performed for the 

BDS and the MBE. The scatter associated with 

plates thinner than 3

/8 inch was greater, leading to less favorable ɸ-factors. Therefore, for 

design, it was assumed that a 3

/8-inch thickness 

limit could be imposed for the gusset plates 

and that the data for “thin” gussets could 

be ignored. However, because the existing 

inventory of gusset plates has “thin” plates 

in it, the statistics used in determining MBE 

ɸ-factors included all the data. 

In design, the resistance factors were 

selected at a DL/LL ratio of 6.0 because this 

represents the most conservative value for 

Strength I because Strength IV will likely control at higher DL/LL ratios. This ratio would 

yield conservative resistance factors for more 

lightly loaded trusses. In addition, design 

resistance factors were calibrated at a target 

reliability index of 4.5. In rating, a different 

philosophy was adopted, primarily because of 

economics. (Using the design philosophy in 

rating would require posting of many trusses 

that have performed satisfactorily.) Resistance 

factors at a reliability index of 3.5 and at a 

DL/LL ratio of 1.0 were selected, but the use of 

system reliability factors for gusset plate rating was mandated. Long span trusses would 

unconservatively be rated because of the 

DL/LL effect, and an additional reduction factor 

was provided to account for the reduction in 

the resistance factor between DL/LL ratios of 

1.0 and 6.0.

Table 1 shows the calibrated resistance factors that were recommended to AASHTO for 

the limit-states discussed in this TechBrief for 

both the BDS and MBE. The resistance factors 

are rounded to the nearest 0.05, in accordance with accepted practice. It should be 

—Not considered.

a

Calibrated at a reliability index of 4.5 at a DL/LL ratio of 6.0.

bCalibrated at a reliability index of 3.5 at a DL/LL ratio of 1.0.

c

Assumes a shear reduction factor, Ω=0.88.

BDS = Bridge Design Specifications.

MBE = Manual for Bridge Evaluation.

Resistance Check BDSa MBEb

Full and partial plane shear yielding 0.83c 1.00c

Whitmore buckling (all data considered) — 0.95

Whitmore buckling (neglecting plates thinner than 3/8 inch) 0.75 —

Block shear 0.80 1.00

Chord splice (all data considered) — 0.85

Chord splice (neglecting plates thinner than 3/8 inch) 0.65 —

Table 1. Summary of resistance factors.
11

noted that the calibrated resistance factors 

between full and partial plane shear yielding 

is the same. The Ω-factor was chosen as 0.88 

because it resulted in resistance factors for 

both the BDS and MBE that were automatically rounded to the nearest 0.05. The chosen 

Ω-factor implies that gusset plates cannot fully 

plastify (as described earlier in this TechBrief); 

this Ω-factor is an artifact of the calibration, 

and its continued use distinguishes between 

shear yielding of gusset plates and regular 

flexural members. 

The resistance factors are different between 

the two specifications because the resistance 

factors were selected at different DL/LL ratios 

and reliability indices for the reasons explained 

above. In particular, the existing BDS resistance factor of 0.80 for block shear was found 

to have a reliability of 4.5. For rating, this resistance factor was increased to 1.00 to reflect the 

rating reliability index of 3.5. 

FUTURE WORK

There are two recommended areas for future 

work not addressed by this project. First, a 

live-load study of trusses should be performed 

to allow the use of less conservative live load 

statistics in the calibration. This study would 

have to consider a wide range of truss span 

lengths from approximately 40 to 800 ft. 

Second, more work should be performed 

to refine the strength calculations of gusset 

plates around tension members. The three 

models that failed in the parametric study 

demonstrated coupled failures of an order 

beyond the simple Whitmore section check. 

The coupled mechanisms appeared to be the 

failure of the gusset on the boundary with the 

chord and vertical, which would be a combined tension and shear failure. The application of the block shear check did not predict 

failure well in these situations. In addition, the

block shear check needs to be validated when 

failure planes can be merged between closely 

spaced tension members, creating a failure 

surface distorted from what is normally used 

in evaluating block shear.

REFERENCES

1. National Transportation Safety Board, 

Collapse of I-35W Highway Bridge: 

Minneapolis, Minnesota, August 1, 2007, 

Highway Accident Report NTSB/HAR08/03, Washington, DC, 2008.

2. U.S. Department of Transportation, LoadCarrying Capacity Considerations of 

Gusset Plates in Non-Load-Path-Redundant 

Steel Truss Bridges, Technical Advisory 

5140.29, Federal Highway Administration, 

Washington, DC, 2008.

3. American Association of State Highway 

and Transportation Officials, AASHTO LRFD 

Bridge Design Specifications–4th Edition, 

Washington, DC, 2007.

4. Ibrahim, F.I.S., Load Rating Guidance and 

Examples for Bolted and Riveted Gusset 

Plates in Truss Bridges, FHWA-IF-09-014, 

Federal Highway Administration, 

Washington, DC, February 2009.

5. American Association of State Highway 

and Transportation Officials, Guide 

Specifications for Strength Design of Truss 

Bridges (Load Factor Design), Washington, 

DC, 1985.

6. Whitmore, R.E., Experimental Investigation 

of Stresses in Gusset Plates, Bulletin 

No. 16, Engineering Experiment Station, 

University of Tennessee, May 1952.

7. Transportation Research Board, Guidelines 

for the Load and Resistance Factor Design 

and Rating of Riveted, Bolted, and Welded 

Gusset-Plate Connections for Steel 

Bridges, NCHRP Project 12-84, National 

Cooperative Highway Research Program, 

Washington, DC, available at http://apps.

trb.org/cmsfeed/TRBNetProjectDisplay.

asp?ProjectID=2509. 
12

AUGUST 2014 FHWA-HRT-14-063

HRDI-40/08-14(150)

Researchers—Justin Ocel of FHWA was the principal investigator for National Cooperative Research Program Project 

12-84. The preceding principal investigators were Joey Hartmann and William Wright, both of FHWA during their 

tenure as principal investigator. The experimental work conducted at Turner-Fairbank was managed by Robert Zobel of 

Professional Service Industries, and digital images correlation data were collected and analyzed by Mark Iadicola of the 

National Institute of Standards and Technology. Georgia Institute of Technology was also a party to the research under 

the direction of Professors Don White and Roberto Leon, with work performed by Yoon Duk Kim, Yavuz Mentes, Towhid 

Bhuiyan, and Joon Wan Hu.

Distribution—This TechBrief is being distributed according to a standard distribution. Direct distribution is being made 

to the Divisions and Resource Center.

Availability—This TechBrief may be obtained from the FHWA Product Distribution Center by email to 

report.center@dot.gov, fax to (814) 239-2156, phone to (814) 239-1160, or online at http://www.fhwa.dot.gov

/research.

Key Words—Gusset plate, buckling, shear yielding, truss bridge, chord splice, Whitmore, bridge design, bridge rating, 

steel bridge.

Notice—This document is disseminated under the sponsorship of the U.S. Department of Transportation in the 

interest of information exchange. The U.S. Government assumes no liability for the use of the information contained 

in this document. The U.S. Government does not endorse products or manufacturers. Trademarks or manufacturers’ 

names appear in this report only because they are considered essential to the objective of the document.

Quality Assurance Statement—The Federal Highway Administration (FHWA) provides high-quality information to 

serve the Government, industry, and public in a manner that promotes public understanding. Standards and policies 

are used to ensure and maximize the quality, objectivity, utility, and integrity of its information. FHWA periodically 

reviews quality issues and adjusts its programs and processes to ensure continuous quality improvement.

8. American Association of State Highway 

and Transportation Officials, AASHTO 

LRFD Bridge Design Specifications—6th 

Edition, Washington, DC, 2012.

9. American Association of State Highway 

and Transportation Officials, The Manual 

for Bridge Evaluation–2nd Edition, 

Washington, DC, 2011.

10. Huns, B. S., Grondin, G. Y., and Driver, R. G. 

“Tension and Shear Block Failure of Bolted 

Gusset Plates,” Canadian Journal of Civil 

Engineering, 33(4), pp. 395-408, April 2006.

11. Kulicki, J.M., Prucz, Z., Clancy, C.M., Mertz, 

D.R., and Nowak, A.S., Updating the 

Calibration Report for AASHTO LRFD Code, 

National Cooperative Highway Research 

Program Report 20-07/186, Transportation 

Research Board, Washington, DC, January 

2007.

12. American Institute of Steel Construction, 

Steel Construction Manual—13th Edition, 

Chicago, IL, 2008.