A Strong Energy Portfolio for a Strong America 

Energy efficiency and clean, renewable energy will mean a stronger economy, a cleaner environment, 

and greater energy independence for America. Working with a wide array of state, community, industry, and 

university partners, the U.S. Department of Energy’s Office of Energy Efficiency and Renewable Energy 

invests in a diverse portfolio of energy technologies.

For more information contact:

EERE Information Center

1-877-EERE-INF (1-877-337-3463)

www.eere.energy.gov

U.S. Department of Energy 

Office of FreedomCAR and Vehicle Technologies 

 1000 Independence Avenue S.W. 

Washington, DC 20585-0121 

FY 2006 

Progress Report for High Strength Weight Reduction 

Materials 

Energy Efficiency and Renewable Energy 

Office of FreedomCAR and Vehicle Technologies 

Advanced Materials Technologies 

Edward Wall Program Manager, OFCVT 

Rogelio Sullivan Advanced Materials Technologies Team Leader 

James Eberhardt Chief Scientist 

March 2006 

High Strength Weight Reduction Materials FY 2006 Progress Report 

CONTENTS

1. INTRODUCTION................................................................................................................................... 1 

2. MATERIALS DEVELOPMENT .......................................................................................................... 3 

A. Advanced Materials for Friction Brakes ............................................................................................ 3 

B. Integrated Approach for Development of Energy-Efficient Steel Components 

for Heavy Vehicle and Transportation Applications.......................................................................... 9

C. Development of Technologies for the Application of Magnesium Metal Matrix 

Composites for Heavy Vehicles ......................................................................................................... 15 

3. MATERIALS PROCESSING TECHNOLOGIES.............................................................................. 25 

A. Development of an Advanced Squeeze Casting Process for Producing High-Integrity

 Truck Components ............................................................................................................................. 25 

B. Counter Gravity (Hitchiner) and Pressure-Assisted Lost Foam Magnesium Casting ........................ 31 

C. Wrought Magnesium Alloy/Process Development ............................................................................ 39 

D. Alloy Design and Thermomechanical Processing of a Beta Titanium Alloy for a Heavy

 Vehicle Application............................................................................................................................ 45 

E. Durability of Titanium Chassis Components............................................................................... 55 

4. ENABLING TECHNOLOGIES............................................................................................................ 59 

A. Joining of Advanced Materials by Plastic Deformation..................................................................... 59 

B. Attachment Techniques for Heavy Truck Composite Chassis Members........................................... 69 

C. Friction-Stir-Joined Aluminum Sheet Materials for Heavy Vehicle Cab Structures ......................... 79 

D. Friction Stir Joining and Processing of Advanced Materials Including MMCs................................. 85 

E. Friction Stir Welding and Processing of Advanced Materials ........................................................... 91 

F. Basic Studies of Ultrasonic Welding for Advanced Transportation Systems ................................... 101 

5. LIGHTWEIGHT VEHICLE STRUCTURES ..................................................................................... 111 

A. Application of Superplastically Formed Aluminum for Truck Body Panels ..................................... 111 

B. Advanced Composite Support Structures........................................................................................... 115 

C. New-Generation Frame for Pickup/Sport Utility Vehicle Application .............................................. 119 

D. Advanced Superplastic Forming Development for Heavy Vehicle Structures .................................. 123 

E. Development of Magnesium for Heavy Vehicle Powertrain Components ........................................ 129 

F. Lightweight Stainless Steel Bus Frame—Phase III............................................................................ 133 

6. APPLICATION OF INNOVATIVE MATERIALS ............................................................................ 139 

A. Application of Carbon Fiber for Large Structural Components......................................................... 139 

B. Hybrid Composite Materials for Weight-Critical Structures.............................................................. 157

C. High-Conductivity Carbon Foam for Thermal Management in Heavy Vehicles............................... 161 

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FY 2006 Progress Report High Strength Weight Reduction Materials 

iv
High Strength Weight Reduction Materials FY 2006 Progress Report 

1. INTRODUCTION 

The energy consumption of on-road heavy trucks has been increasing because of the growth in demand 

for freight delivery as the U.S. economy continues to expand. The recent rapid rise in fuel prices has elevated 

fuel costs to the number one concern of the trucking industry. Owners and operators of trucks and truck fleets 

have been operating on razor-thin margins, and increasing the fuel economy of their vehicles is one way of 

reducing their fuel costs to remain viable as a business. One way of achieving heavy vehicle fuel economy 

improvements, without sacrificing revenue earnings limited by maximum allowable vehicle weight or 

volume, is through the use of high-strength weight reduction materials for the tractor or trailer or both. 

Vehicle weight reduction results in reduced parasitic energy losses and improved overall vehicle fuel 

efficiency by enabling more ton-miles per gallon of fuel consumed. In addition, it is recognized that improved 

materials may enable the implementation of other technologies that can further improve the fuel efficiency of 

heavy vehicles. 

The High-Strength Weight Reduction (HSWR) Materials Technology activity of the U.S. Department of 

Energy’s (DOE’s) Office of FreedomCAR and Vehicle Technologies (OFCVT) Program seeks to reduce the 

weight of heavy vehicles without reducing vehicle functionality, durability, reliability, or safety and to do so 

cost-effectively. Work has been focused on developing advanced materials and materials processing 

technologies that can be applied to a wide array of heavy vehicle body, chassis, and suspension components 

to achieve weight reduction. Research is focused on overcoming barriers to the widespread introduction of 

lightweight materials in the heavy vehicle industry. Major barriers include cost; design and simulation 

technologies; manufacturability; prototyping and tooling; joining and assembly; and maintenance, repair, 

durability, and recycling. Priority materials include advanced high-strength steels, aluminum, magnesium, 

titanium, and composites such as metal matrix materials and glass- and carbon-fiber-reinforced thermosets 

and thermoplastics. The research required to develop these technologies is too high-risk to be pursued 

independently by the heavy vehicle industry because of substantial return-on-investment uncertainties. In 

addition, the relatively low volume of truck components makes it difficult for any individual truck 

manufacturer to develop on its own a new production/fabrication process using new kinds of materials such 

as composites. 

Research and development (R&D) activities were conducted through a variety of contractual 

mechanisms. These include cooperative research and development agreements (CRADAs), university grants, 

R&D subcontracts, and directed research. Research partners include heavy vehicle manufacturers (including 

member companies of the government/industry 21st Century Truck Partnership), first-tier component 

manufacturers and materials suppliers, national laboratories, and other non-profit technology organizations. 

Laboratories include Argonne National Laboratory (ANL), Idaho National Engineering and Environmental 

Laboratory (INEEL), Los Alamos National Laboratory (LANL), Oak Ridge National Laboratory (ORNL), 

and Pacific Northwest National Laboratory (PNNL). 

Among the several goals established for this development area, one is to enable at least a 22% reduction 

in the weight of an unloaded tractor-trailer combination and about a 10–33% reduction in the weight of other 

classes of heavy vehicles, depending on performance requirements and duty cycles. With the completion and 

issuance of this final High Strength Weight Reduction Materials report, most of these goals have been 

accomplished. Technology R&D results have been made available to industry for further development and 

commercialization. Some of these results are discussed in this final annual report. 

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FY 2006 Progress Report High Strength Weight Reduction Materials 

2 
High Strength Weight Reduction Materials FY 2006 Progress Report 

2. MATERIALS DEVELOPMENT 

A. Advanced Materials for Friction Brakes 

Principal Investigator: P. J. Blau 

Oak Ridge National Laboratory 

P.O. Box 2008, Oak Ridge, TN 37831-6063 

(865) 574-5377; fax: (865) 574-6918; e-mail:blaupj@ornl.gov 

Technology Development Area Specialist: James J. Eberhardt 

(202) 586-9837; fax: (202) 587-2476; e-mail: james.eberhardt@ee.doe.gov 

Field Technical Manager: Philip S. Sklad 

(865) 574-5069; fax: (865) 574-6098; e-mail: skladps@ornl.gov 

Contractor: Oak Ridge National Laboratory 

Contract No.: DE-AC05-00OR22725 

Objectives 

• Determine the feasibility of using advanced, lightweight materials and surface treatments for truck disk brake 

surfaces and develop the friction, wear, and performance data needed to promote their introduction. Materials 

of interest include titanium alloys and their composites, carbon-based materials, and novel surface treatments 

and coatings that enhance both durability and frictional stability over a range of temperatures. 

Approach 

• Design and build a subscale brake material testing apparatus to conduct friction and frictional-heating studies 

of advanced materials at speeds and pressures similar to those of full-sized truck brakes. 

• Characterize changes to the surfaces of materials that occur as a result of frictional contact under braking conditions. Relate these changes to material properties and frictional performance. 

• Investigate the effects of speed, contact pressure, and temperature on the frictional response of candidate brake 

rotor materials, such as titanium alloys, thermally sprayed coatings, ceramic composites, and carbon materials, 

and apply that knowledge to identify which of them seem most promising for further development. 

• Identify and test friction materials (pads) that work well with the advanced brake rotor materials. 

• Work with trucking industry partners and commercial dynamometer test laboratories to evaluate advanced candidate materials under established testing conditions, such as standard American equivalent (SAE) procedures 

and federal test methods. 

Accomplishments 

• Designed and built an instrumented, subscale brake materials friction and wear testing system. 

• Investigated the effects of wet and dry braking conditions on frictional behavior. 

• Published a review of the compositions of friction materials that includes thermo-physical properties data. 

• Characterized wear particles that are generated in tests of candidate advanced brake materials. 

• Published a study of the effects of the wear test method on relative rankings of commercial brake linings. 

3 
FY 2006 Progress Report High Strength Weight Reduction Materials 

• Published a technical report that summarizes current ORNL research on nontraditional brake materials, such as 

ceramic composites, titanium alloys, and titanium composites. 

• Friction-tested nontraditional brake materials on a titanium (Ti)-based composite that was produced using lowcost Ti powder and compared the results to tests using a novel thermal diffusion treatment to harden the surface 

of a traditional ingot metallurgy Ti alloy. 

• Worked with a commercial truck industry supplier to test thermally sprayed Ti rotors in a full-scale dynamometer and captured thermal images to compare their heating response to traditional cast iron. 

• Prepared new carbon-based brake composites for friction testing on the subscale brake tester (SSBT). 

Future Direction 

• Complete industry testing protocols on full-scale, thermally sprayed coated titanium truck brake rotors. 

• Complete tests on densified carbon-based disk materials and on a novel surface treatment for Ti-based rotors. 

Introduction 

New, aerodynamic designs for energy-efficient 

heavy trucks and advanced tires with lower rolling 

resistance can markedly improve fuel economy. 

However, as parasitic losses and the drag on trucks 

decrease, the demands on brakes increase. Braking 

systems include hardware design, instrumentation 

and control, and friction materials that dissipate energy as the vehicle decelerates. This project specifically addresses the latter. Friction brake materials 

must exhibit a combination of characteristics, including 

• frictional stability over a wide range of temperatures 

• good frictional heat removal and dissipation 

• dimensional stability to avoid vibrations 

• corrosion resistance to road de-icing compounds 

• durability and wear resistance 

• lightweight and cost competitiveness 

Opportunities exist to utilize advanced materials 

to improve the fuel efficiency without compromising 

vehicle safety and reliability. 

The goal of this project is to characterize newly 

developed materials and surface treatments that 

show significant potential as truck brake rotor materials. The approach is first to characterize and friction test small specimens of candidate materials in 

the laboratory and then progress to full-scale component tests using standard industry procedures. For 

laboratory testing, an instrumented, SSBT was built 

and used to evaluate a wide variety of advanced materials, including ceramic composites, intermetallic 

alloys, titanium alloys, metal matrix composites, and 

hard-coated titanium. A summary of the test results 

was prepared and distributed to U.S. truck manufacturers and braking system component suppliers. 

Friction and wear testing is coupled with microstructural studies of surface films generated during 

braking and correlated with measurements of physical properties. A thermal efficiency parameter that 

expresses the rise in surface temperature per unit of 

frictional work was developed to assess the extent to 

which brake design features, such as vents, must be 

taken into account when replacing traditional cast 

iron rotors. During this reporting period, a full-scale, 

titanium alloy truck brake rotor was tested for the 

first time in a commercial truck component manufacturer’s facility (Figure 1). Results were particularly promising in terms of the resistance of the surfaces to fade at high temperatures. 

Materials of Current Interest 

A number of candidate materials were tested and 

then eliminated from further consideration due either 

to their low melting points, brittleness, sensitivity to 

corrosion, or wear-particle characteristics. Those of 

current interest are shown in Table 1. 

4 
High Strength Weight Reduction Materials FY 2006 Progress Report 

Figure 1. One of the first four full-scale, thermally 

sprayed titanium truck brake rotors to be produced. It 

was prepared for this project by Red Devil Brakes, 

Inc., of Pennsylvania. 

Table 1. Candidate materials of current interest for friction studies on the SSBT apparatus 

Rotor materials Source 

Titanium alloy, thermally sprayed with a 

proprietary frictioncontrol coating 

Titanium powder composite with hard particles in the surface layers 

Oxygen diffusiontreated titanium alloy 

Densified carbon material 

Red Devil Brakes, Inc., Mount 

Pleasant, Pennsylvania 

C. Blue and W. Peter, ORNL, 

in collaboration with International Titanium Powder Corporation 

A process being developed for 

improving the tribology of 

titanium surfaces by J. Qu, 

ORNL 

Developed by J. Klett, ORNL 

The thermally sprayed and powder metallurgy 

Ti materials have been described in earlier reports. 

The coating consists of a proprietary blend of metallic and ceramic particles. The substrate is Ti-6Al-4V 

alloy. 

The densified carbon materials were prepared 

during this reporting period at ORNL. Plates were 

produced in two forms: a two-dimensional (2-D) 

compressed structure (Figure 2) and a threedimensional (3-D) weave (Figure 3). One of the 

principal differences in these structures is in the heat 

flow. The 3-D material tends to remove heat from 

the surface somewhat better, but the 2-D structure 

has better in-plane conduction. After preparation, 

they were machined to fit the mounting hardware in 

the SSBT. The initial selection of counterface (pad) 

materials for the carbon disk specimens was based 

on prior experience in testing C-SiC composite disks 

in a project with GE Power Systems (formerly Honeywell Advanced Composites, Inc.). In that work, it 

was found that the best performance involved testing 

with self-mated materials. However, a commercial, 

semimetallic pad material may also be tried, depending on the results of the first friction test trials on the 

SSBT. 

Figure 2. Surface of the 2-D structure, as finished. 

Figure 3. Surface of the 3-D structure, as finished. 

Dynamometer Tests and Thermal 

Imaging 

One of the four thermally sprayed Ti rotors 

(shown in Figure 1) was shipped to a major truck 

5 
FY 2006 Progress Report High Strength Weight Reduction Materials 

component manufacturer for dynamometer testing in Rotor surface temperature after 4 seconds 

comparison with a standard cast iron rotor of the during 50-15 mph snubs 

1200 

same type. J. Martino, of Red Devil Brakes, prepared the brake pads for that test. In addition to a 1000 

Gray Cast Iron 

Th Spr Titanium 

T max = 277.8 + 76.66 N - 3.968 N2 

T max = 706.15 + 96.43 N - 5.455 N2

modified version of a federal testing protocol, a series of rapid decelerations (“snubs”) from equivalent 

highway speeds of 50 mph down to 15 mph, were 

performed. R. Dinwiddie, of ORNL, was present to 

800 

T (

oC) max

600 record calibrated thermal images of the cast iron rotor and the Ti rotor using a high-speed, rotationsynchronized infrared camera. 400 

A thermal image of the Ti rotor after a series of 

five snubs is shown in Figure 4. The maximum heat- 200 

ing seemed to occur about four seconds after the 

brake pads were applied. Figure 5 compares these 0 

maximum temperatures for the Ti and cast iron 

discs. 

Figure 4. Infrared (IR) thermal image of the Ti rotor 

during the sixth snub stop. Red represents a temperature of about 1600oF and dark blue about 1200oF. 

Dashed dark streaks at the right edge of the disk are 

cooling vents. 

As Figure 5 data show, the Ti disc ran several 

hundred degrees hotter than the cast iron disk, but it 

retained its frictional capabilities at these higher 

temperatures, an important consideration in emergency stops, repeated brake application during long 

hills, and in the kind of service that requires frequent 

stops, such as school buses, delivery trucks, and 

trash haulers. 

0 1 2 3 4 5 6 7 

SNUB NUMBER (N) 

Figure 5. Friction coefficients (μave) after four seconds of 

sliding for cast iron and Ti brake rotors. 

Oxygen Diffusion Treatment Appears 

Promising 

In addition to thermal spraying and the addition 

of hard-particles, a novel process called thermal diffusion has been used to improve the frictional characteristics of Ti-6Al-4V alloy surfaces. Preliminary 

SSBT experiments, conducted during this period, 

show excellent frictional behavior over a range of 

testing speeds and temperatures. A patent disclosure 

has been filed and details will be reported later. 

Future Plans. Testing of carbon-based brake disc 

materials. SSBT friction tests of the two forms of 

densified carbon described earlier in this report will 

be performed in the fall of 2006 and results are expected to be available by the end of CY 2006. Additional analysis of the test data for thermal oxidation 

will be performed, and a paper comparing it to alternative treatments is planned. 

Conference Paper. A summary of the past two 

years of research on Ti-based materials for disc 

brakes has been submitted to the 16th International 

Conference on Wear of Materials (2007) . Plans for 

the conference include a special session on brake 

materials and that will be an excellent forum for the 

presentation of the current results. 

Conclusions 

The friction and frictional-heating response of 

several Ti-based brake rotor materials have been 

6 
High Strength Weight Reduction Materials 

evaluated. The materials’ excellent retention of 

stopping capabilities at elevated temperature (“fade 

resistance”) was demonstrated both in the laboratory 

and in full-scale tests of a thermally sprayed truck 

brake rotor. 

Publications/Presentations/Patents 

Peter J. Blau, Brian C. Jolly, William H. Peter, 

and Craig A. Blue, “Tribiological Investigation of 

FY 2006 Progress Report 

Titanium-Based Materials for Brakes,” accepted for 

Wear of Materials ‘07, Montreal, Canada, April 15– 

19, 2007. 

Patent disclosure: “Oxygen-Diffused Titanium 

for Engine and Other Bearing Applications.” 

7 
High Strength Weight Reduction Materials FY 2006 Progress Report 

B. Integrated Approach for Development of Energy-Efficient Steel Components 

for Heavy Vehicle and Transportation Applications 

Co-Principal Investigator: Leo Chuzhoy 

Caterpillar, Inc. 

P.O. Box 1875, Peoria, IL 61656-1875 

(309) 578-6621; fax: (309) 578-2953; e-mail: chuzhoy_leo@cat.com 

Co-Principal Investigator: Gerard Ludtka 

Oak Ridge National Laboratory 

P.O. Box 2008, Oak Ridge, TN 37831-6064 

(865) 574-5098; fax: (865) 574-3940; e-mail: ludtkagm1@ornl.gov 

Co-Principal Investigator: Alan Needleman 

Brown University 

Box D, Brown University, Providence, RI 02912-D 

(401) 863-2863; fax: 401-863-1157; e-mail: alan_needleman@brown.edu 

Chief Scientist: James J. Eberhardt 

(202) 586-9837; fax: (202) 587-2476; e-mail: james.eberhardt@ee.doe.gov 

Field Technical Manager: Philip S. Sklad 

(865) 574-5069; fax:(865) 574-6098; e-mail:skladps@ornl.gov 

Contractor: Caterpillar Corp., ORNL, Brown University 

Contract No.: DE-AC05-00OR22726, DE-AC05-00OR22725, 4000023957

Objective 

• Develop microstructure-level (often called mesoscale) simulation (MLS) tools to capture the formation and influence of inhomogeneous (real life) microstructures in steel processing. 

• Use the tools to understand and predict microstructure evolution during processing and ultimately to predict the 

resultant component performance. 

• Use the information gained to design steel microstructures and develop process roadmaps. 

• Demonstrate the developed techniques on a pilot project at Caterpillar involving a defined structural component, 

a track roller shaft, that represents a steel component with high production volume that could benefit from microstructure-level improvements. 

Approach 

• Develop microstructure-level models to accurately predict the evolution and behavior of steel microstructures 

during component processing and performance. 

• Characterize thermomechanical properties of microstructural constituents as a function of temperature and composition for exact chemistries. 

• Develop tools to assess required environmental resources and integrate them into the simulation. 

9 
FY 2006 Progress Report High Strength Weight Reduction Materials 

Accomplishments 

• Developed and validated specific MLS models for at least one experimental case: 

— Developed the first reported quenching model for the transformation of austenite to lath martensite. 

— Obtained predicted transformation rates in general agreement with experimental results. 

— Used the model to show the dynamic microstructure evolution during quenching and the effect of process 

variables on microstructure evolution. 

— Developed a microstructure solidification model to predict MnS inclusion forming during solidification. 

— Developed a damage evolution model at the microstructure level to predict damage initiation and cyclic fatigue failure. 

— Developed a microstructure evolution model during induction heating to predict grain size and austenitization fraction. 

Future Direction 

• Fully couple the phase transformation and evolution prediction of the heat treatment module with mechanical 

property evolution and stress state evolution. 

• Build links to transfer the microstructure information combined with mechanical properties and stress into the 

microstructure-level failure model. 

• Develop a microstructure-level model that fully couples phase transformation, mechanical property evolution, 

and residual stress evolution. 

• Establish an integrated method to take microstructure information from the manufacturing process into failure 

analysis. 

Introduction 

Product cost and performance are two major 

pressures influencing the acceptance of energy savings technologies by manufacturers, suppliers, and 

users in the heavy vehicle and transportation industries. Energy cost has become a significant portion 

of total product cost for steel applications. Major 

reductions in energy use are potentially achievable 

for the transportation and heavy vehicle industries 

through development and optimization of costeffective fabrication processes and enhanced product 

performance. The key enabling technologies to 

achieve these benefits are improved materials and 

realistic, microstructure-level simulations to predict 

manufacturability and life-cycle performance. Over 

the past two decades, steel mills and forge shops 

have successfully implemented numerous energyefficient processes. The next logical step is to focus 

on the development of steel microstructures produced in such a way that they are energy-efficient 

and environmentally benign over the entire manufacturing cycle. 

Structural materials used in critical steel components of machines have evolved to the point where 

further improvements in the performance can be 

achieved only through a fundamental understanding 

of the mechanisms driving material behavior during 

processing and service. Microstructural elements 

such as grain size, inclusion and precipitate distributions, and chemistry control the performance of engineering materials. Variation in these 

microstructural elements leads to variation in such 

critical properties as fatigue life, toughness, and 

wear resistance. Therefore, understanding and developing the capability to control the formation of 

steel microstructures and predicting the functional 

and environmental performance is critical to moving 

the industry closer to its energy-efficiency, resourceefficiency, and pollution-prevention goals. The full 

realization of these benefits, however, requires a 

design tool that optimizes the microstructure with 

respect to the mechanical and environmental performance throughout the life cycle of a particular 

steel component. 

To apply an integrated approach to the development of energy-efficient steel components, the 

development of three major areas of research needs 

to be completed: 

• microstructure-level models to accurately predict the evolution and behavior of steel microstructures during component processing and 

performance 

10 
High Strength Weight Reduction Materials 

• thermomechanical properties of microstructural 

constituents as a function of temperature and 

composition for exact chemistries 

• tools to assess required environmental resources 

This project addresses activities required for the 

first of these three areas. The second and third tasks 

are either supported by current funding through the 

National Science Foundation or are expected to be 

funded through the DOE Initiative for Proliferation 

Prevention. Through these integrated efforts, a design tool will be developed that optimizes the microstructure, manufacturability, and performance of 

components with respect to the mechanical and environmental performance required throughout the 

components’ life cycles. The overall benefit of this 

research will be the development and demonstration 

of a design methodology that will enable the domestic transportation and heavy vehicle industries to 

compete effectively in future worldwide markets 

through improved product performance and energy 

savings. Furthermore, the proposed design tool can 

be extended to other materials—i.e., cast iron, aluminum, titanium, magnesium, nickel-based alloys, 

ceramics, and composites—thus impacting virtually 

all industries. 

In addition to significant energy and cost savings 

benefits to the transportation and steel industries 

through microstructure-level modeling, virtually all 

industries will be impacted. Examples include aerospace (engines, transmissions, structural), marine 

(engines and drives), agricultural and construction 

equipment, oil and chemical processing (pumps and 

gear boxes), military vehicles, mining machinery, 

appliances (compressors, motors, gear boxes, 

shafts), power tools, and automotive aftermarket. 

Project-end Deliverables 

The deliverables for the project will be the following: 

• Microstructure-level methods for simulating the 

manufacturing cycle of steel products coupled 

with material performance computations (Tasks 

1, 2, and 3) 

• A specific chemical composition and processing 

map for a 1500-series steel and a micro-alloyed 

steel suitable for the application under consideration (Task 4) 

FY 2006 Progress Report 

• Energy and environmental resources required to 

produce the selected steel component using at 

least two steels and processing schemes (Task 4) 

The project tasks will focus on three critical areas in the development of these models: heat treatment processing, machining, and materials 

performance in specific applications. These three 

processes have been chosen because they are critical 

steps in reaching the goal of being able to develop a 

specific steel for a particular application with a 

heavy reliance on modeling and with energy requirements optimized as an integral part of the development process. Modeling of casting and forming 

processing has been performed under other programs, and Caterpillar has shown successful application and commercialization of these models. These 

existing models will be used along with the heat 

treatment, machining, and specific applications 

models described in this report to complete a suite of 

models for the manufacture of micro-alloyed steels. 

To accomplish the project objectives, a multidisciplinary team consisting of a national laboratory, 

a university, and a steel end user has been assembled. The team will be supported by an international 

research institute for material characterization and 

by experts in environmental impact assessment. The 

following section describes each of the technical 

tasks to be performed. 

Semiannual Period Report Topics 

Model development and results for the following two topics were reported in detail in the 2006 

semiannual report and therefore will not be duplicated in this annual report: 

• The first reported quenching model for the 

transformation of austenite to lath martensite has 

been developed. 

• The predicted transformation rates are in general 

agreement with experimental results. 

• The model can show the dynamic microstructure 

evolution during quenching and the effect of 

process variables on microstructure evolution. 

11 
FY 2006 Progress Report 

Modeling of Fatigue Life Prediction Based on 

Continuum Damage Mechanics 

Background 

Fracture and fatigue failures are always important issues to be considered by the engineers in designing engineering components. About 80–90% of 

all engineering failures relate to fracture and fatigue 

for different length scales and different material systems 

Investigations of failure and crack growth related to fatigue have been a challenge to scientists 

for decades. Classical linear elastic fracture mechanics has been successful in dealing with fracture problems with a negligible nonlinear zone ahead of the 

crack tip. However, the concepts of fracture mechanics present difficulty when the nonlinear process 

zone is relatively large compared with the crack 

length or the size of the structure, or on a microstructural length scale. 

Void nucleation and growth has been recognized 

as a key micromechanism of rupture for ductile metals. In a continuum damage mechanics (CDM) approach, damage is assumed to be one of the internal 

constitutive variables that account for the effects on 

the material associated with a given damage condition through the definition of a state variable of 

damage. Damage not only affects the material’s 

yield function but also reduces the stiffness through 

the equation of effective stresses with cyclic loading. 

In addition, the fatigue properties of metals are quite 

microstructure sensitive. The CDM approach is 

promising for investigating the relationship between 

the fatigue damage properties and microstructure. 

Objective 

The goal of the project is to develop a microstructure-level material simulation model based on 

metal plasticity and CDM by accounting for the material degradation during cyclic loading to predict 

the fatigue failure and service life of components. 

The material properties are related to the microstructure of the material. An MLS is developed to consider the void nucleation stage from inclusions and 

porosity through void growth and coalescence and 

ultimately to the final material failure stage. This 

simulation technique clearly addresses the link between macro-level failure and microstructure-level 

defects. Microstructure changes during the manufacturing process then can be integrated into the failure 

High Strength Weight Reduction Materials 

prediction further. A more accurate life prediction 

based on developed CDM based modeling is 

achieved. 

The occurrence of plastic deformation and accumulation with loading cycles is a main cause of 

damage development that will lead the material to 

fail later. Under cyclic loading, irreversible microstructure changes occur in the following forms: persistent slip bands, rearrangement of dislocation 

systems into cell structures, and void nucleation and 

growth at secondary phase inclusions. The latter is 

one cause of the high strain, which is expected to 

result in a short life. 

The major advantage of the approaches based on 

continuum mechanics is the transferability from 

specimen to structure, since model parameters 

should be geometry-independent. 

Methodology 

This model development consisted of the following steps: 

• Development of a damage evolution law for 

steel under cyclic plastic flows 

• Development of the approach to extract the material parameter from the experimental data 

• Design of a program to simulate the damage 

evolution during the cyclic loading and a damage-driven fatigue model 

• Validate the model and the accuracy of the 

simulation results by comparing them with the 

experimental S-N curve 

Results 

The described methodology was used to characterize the material parameter in the derived damage 

evolution fatigue model (specifics will be published 

in a subsequent report) for a gray iron material (class 

40 B). A typical testing curve is shown in Figure 1. 

To clearly see the loading-unloading curve for different hysteresis loops during multiple fatigue cycles, successive curves are offset by 0.005 strain in 

Figure 2. The test was conducted at room temperature and was controlled by a constant strain amplitude condition. 

In Figure 3, the elastic modulus regression is 

plotted with the increasing number of cycles. It can 

be seen clearly that the modulus under the tension 

portion of the curve keeps dropping with increasing 

12 
High Strength Weight Reduction Materials 

Figure 1. Schematic drawing to calculate the increment of the plastic strain. 

number of cycles, which means more damage accumulation. Also, the modulus regression data show 

that the effect of damage to modulus regression under compression is not as much as the effect under 

tension. 

200 

150 

100 

50 

0 

-50 

-100 

-150 

-200 

Stress,

FY 2006 Progress Report 

The corresponding damage evolution with number of cycles is shown in Figure 4. It can be seen 

that at the beginning, the damage accumulates 

slowly. With an increasing number of cycles, more 

damage accumulated; and after the damage reached 

the critical damage value Dc=0.3, the damage 

growth rate increased suddenly and material rupture 

occurred. 

Verification

Using the material parameters obtained as described above, the damage model was used to calculate the fatigue life for a different loading range and 

compare it with the experimental S-N curve. Both 

the calculation results and the experimental results 

were plotted on the same graph, Figure 5. The fatigue cycles ranged from 100 to 100000 cycles corresponding to different strain levels. The calculated 

data agree well with the experimental S-N curve. 

This also proves that the CDM approach is applicable to different load levels and fatigue life ranges. 

Note: Successivecurves are offsetby 0.0005 strain. -250 

-0.0050 0.0050 0.0150 0.0250 

Strain 

Figure 2. Fatigue test stress strain curve. 

13 
FY 2006 Progress Report High Strength Weight Reduction Materials 

Figure 3. Elastic modulus regression during the cyclic 

loading condition. 

Figure 4. Damage evolution for the fatigue test. 

Conclusion 

A material model based on CDM was developed. In this model, the damage is coupled with 

plastic strain and provides more accuracy to the 

damage evolution. An approach to calibrate the material parameter was developed for this model. The 

material damage properties (such as initial damage, 

critical damage, and crack closure factor) were extracted from the cyclic fatigue loading test data for 

gray iron. Then, in the simulation, the fatigue model 

was used with the measured damage properties to 

estimate the fatigue life by considering the cyclic 

damage accumulation under cyclic fatigue loading. 

Figure 5. The comparison of experimental and calculated 

S-N curve. 

The life prediction from the simulation shows 

very good agreement with the experimental results. 

The predicted S-N curves (ranging from 100 to 

100000 cycles) agree with the experimental data. 

This proved that this model is robust at different 

load levels as well as in different ranges of cycles in 

fatigue life. This model can be applied to different 

types of metals and materials based on the damage 

evolution constitution. 

Presentations and Publications 

Adrian V. Catalina, Doru M. Stefanescu, Leo 

Chuzhoy, and Michael L. Johnson, “A Solid Fraction Evolution Approach for Modeling of Dendritic 

Growth in Multicomponent Alloys,” presented at the 

TMS Annual Meeting, San Antonio, March 13–16, 

2006. 

Don Sherman, “Characterization of the Carbon 

and Retained Austenite Distributions in Martensitic 

Medium Carbon, High Silicon Steel,” submitted for 

publication in Metallurgical and Materials Transactions A in October 2006. 

B. J. Yang, Leo Chuzhoy, and Mike Johnson, 

“Modeling of Reaustenitization of Hypoeutectoid 

Fe-C Steels with a Cellular Automaton Method,” 

presented at the TMS Annual Meeting, San Antonio, 

March 13–16, 2006. 

B. J. Yang, Leo Chuzhoy, and Mike Johnson, 

“Modeling of Reaustenitization of Hypoeutectoid 

Steels with Cellular Automaton method,” submitted 

to Journal of Acta Mater. in August 2006. 

14 
High Strength Weight Reduction Materials FY 2006 Progress Report 

C. Development of Technologies for the Application of Magnesium 

Metal Matrix Composites for Heavy Vehicles 

Principal Investigator: Adam R. Loukus 

GS Engineering, Inc. 

101 West Lakeshore Drive 

(906) 482-1235; fax: (906) 482-1236; e-mail; adam.loukus@gsengineering.com 

Chief Scientist: James J. Eberhardt 

(202) 586-9837; fax: (202) 587-2476; e-mail: james.eberhardt@ee.doe.gov 

Field Technical Manager: Philip S. Sklad 

(865) 574-5069; fax:(865) 574-6098; e-mail:skladps@ornl.gov 

Participants 

Glen Simula, Andrew Halonen, Tom Wood, GS Engineering, Inc. 

Josh Loukus, Gary Pennala, REL Machine, Inc. 

Robert Hathaway, Chad Johnson, Vadim Pikhovich, Oshkosh Truck Corp. 

Dave Weiss, Eck Industries 

Bob Coleman, Engineered Fiber Solutions 

Contractor: GS Engineering, Inc. 

Contract No.: WR10637 

Objective 

• Analyze different casting processes before designing and building process-controlled squeeze casting equipment. Use this equipment to cast metal matrix composite (MMC) components with mechanical properties adequate to replace heavier materials used in Class 8 vehicles. 

Approach 

• Infiltrate several preforms with the magnesium (Mg) alloy AZ91 using different casting processes: pressureless 

infiltration (PLI), squeeze casting, and pressure infiltration vacuum-assisted casting (PIVAC). 

• Evaluate the microstructural features in the composite structures and evaluate the interactions of the Mg alloy 

and the reinforcement. Follow a parallel path of process simulation to optimize material flow into preforms. 

• Use MAGMAsoft® simulations to assist in the die design, machine a casting die, and make preliminary castings. 

• Pursue further Mg components casting. 

Accomplishments 

• Analyzed the differences in material characteristics, such as the strength, stiffness, grain size, and interfacial reactions, as functions of casting processes. 

• Designed and simulated mold filling with a preform in a connecting rod component. 

• Built a casting die and performed initial infiltration trials on a GSE press with aluminum alloys to streamline 

the casting system. 

15 
FY 2006 Progress Report High Strength Weight Reduction Materials 

Future Direction 

• Milestone 1—Cast magnesium MMC connecting rods. 

• Milestone 2—Evaluate components in bench and field tests. 

Introduction 

The microstructural features of the Mg composites castings done in year 1 have been studied in 

some depth in the past few months. Significant understanding of the interfaces has been developed in 

the past months that has reemphasized the goals and 

route that have been chosen. A clean, inert environment during casting is crucial for the matrix-toreinforcement bonding and hence a high-integrity 

casting. The GSE squeeze casting press is operational, and casting trials are scheduled as the materials arrive to determine the effects of the casting parameters on component strength, stiffness, and porosity. MAGMAsoft software was used to aid in 

process and die design to minimize the chances of 

porosity in the cast components. A detailed plan for 

the remainder of year 2 funds is included in the future work section. 

Grain Structure Analysis 

The squeeze cast trials were run on a production 

caster on which monitoring and control of parameters necessary for consistent quality were not available. PLI studies were performed in order to clarify 

the infiltration potential of various preform materials. Using the PLI process also provided evidence of 

the effects of infiltration velocity and cooling rate on 

microstructure. The controllable casting atmosphere 

in the process eliminates rapid oxide formation that 

is prevalent with the squeeze casting process. 

Metallographic analysis shows the advantages of 

squeeze casting over a PLI process. The cycle time 

for squeeze casting is shorter; hence the cooling rate 

will be higher than for a PLI process. This differential in cooling rate and applied pressure is illustrated 

in the corresponding metallographic structures (Figures 1 and 2). 

The finer grain size achievable with the squeeze 

casting process (Figure 1) should result in greater 

ductility, toughness, and strength compared with the 

PLI process (Figure 2) where the grain size is 3 

times larger. 

Figure 1. Carbon fiber 25% (compacted), AZ91D-T6 squeeze cast (SC) 

process, 100×. Etched with 10% citric 

acid. ASTM E112 grain size 7. 

Figure 2. Carbon fiber 25%, AZ91D-as 

cast, PLI process 100×. Etched with 

10% citric acid. ASTM E112 grain size 

2. 

SEM/Auger Analysis 

The intent of the following analysis was to examine the carbon fiber in two orientations so as to 

identify the reaction products at the interface. Using 

focused ion beam (FIB), Mg from the matrix surrounding the fiber was removed. A sputter profile 

was used to establish how much sputter cleaning 

needed to be done to remove the contaminant and 

native oxide layers formed on the polished surface. 

The data indicated that at least 2 min of sputtering at an etch rate of ~200 nm/min is required. To 

create a smooth sample surface, it was necessary to 

argon-ion sputter while rotating the specimen for 

approximately 3 min. 

16 
High Strength Weight Reduction Materials 

Figure 3 is an scanning electron microscopy 

(SEM) image of an area of the sample that was analyzed to identify reactions along the length of an 

individual carbon fiber. Table 1 shows the atomic 

concentration of the areas analyzed. as indicated in 

Figure 3. All values are relative, not absolute, since 

no standards were used. The spectral data were taken 

after mapping; i.e. the surface had time to adsorb a 

lot of residual gas (e.g., water, CO, CO2). In fact, a 

comparison of points 3 and 5 shows that there are 

nearly equal amounts of oxygen, although area 3 is 

clearly a reaction product. The material in area 3 is 

an oxide reaction product while the oxygen seen in 

area 5 is adsorbed oxygen on the Mg-matrix surface. 

Table 1. Atomic concentration survey spectra 

from the points shown in Figure 3 and resulting compositions 

Are C O Mg Al 

a 

1 4.78 25.07 25.71 44.43 

2 14.28 27.20 35.01 23.51 

3 5.10 34.44 55.64 4.81 

4 73.11 4.51 15.98 6.40 

5 5.02 31.28 59.43 4.27 

6 8.68 30.75 43.08 17.49 

Figure 3. SEM image of a longitudinal 

view of fibers in the matrix. The indicated points are where spectral analysis 

was performed. 

Figure 4 is an on-end SEM image of an area of 

the sample that was analyzed to identify reactions 

around the circumference of an individual carbon 

fiber. Table 2 shows the atomic concentration levels 

from the spectral analysis of the areas identified in 

Figure 4. The analysis shows that the reaction products have high levels of oxygen. 

FY 2006 Progress Report 

Figure 3. SEM image of a longitudinal 

view of fibers in the matrix. The indicated 

points are where spectral analysis was performed. 

Table 2. Atomic concentration of the indicated points from Figure 4 

Area C O Mg Al 

1 2.19 5.46 89.85 2.49 

2 3.95 9.86 67.21 18.98 

3 7.07 27.26 24.14 41.53 

4 75.96 3.16 16.05 4.83 

5 9.26 31.85 56.86 2.03 

The use of auger electron spectroscopy (AES) 

elemental mapping highlights the composition of the 

interfacial constituents. Figure 5 shows that oxiderich areas surround the fiber. Together, the two 

maps show that Mg-O and Al-O reaction products 

are formed at the fiber/matrix interface. 

Figure 5. Individual gray-scale AES maps of Mg, 

carbon, aluminum, and oxygen. These maps correspond to the SEM image in Figure 3. 

17 
FY 2006 Progress Report 

At high magnification, the reaction zone at the 

interface is observed. This zone is critical to the ultimate characteristics of the engineering component. 

In Figure 6, we see the interface containing silicates 

and oxides that can prove detrimental to the material’s mechanical properties. 

Mg Matrix 

Carbon Fiber 

Figure 6. SEM showing the interfacial reaction between the carbon fiber and the magnesium matrix with a sample produced by 

PLI. 

From the AES analysis, it is apparent that essentially four main elements are present in the interfacial regions. Magnesium, oxygen, aluminum and 

silicon have all been mapped as shown in Figure 7. 

It is then of practical and scientific use to determine 

which compounds are present in the casting, as we 

know the constituents. Therefore, some samples 

were scanned using X-ray diffraction to conduct this 

analysis. 

To positively identify secondary phases, the 

samples as identified in Table 3 below were analyzed using X-ray diffraction. Samples were metallographically prepared and etched with 3% nital to 

minimize cold work effects due to polishing. The 

analysis was accomplished using a Scintag XDS2000 θ/θ diffractometer. The X-ray diffraction results confirm the presence of magnesium oxide 

(MgO) and magnesium silicide (Mg2Si) at the fiber/matrix interface that was observed in the AES 

analysis of the interface region in the PLI sample. 

Mold Flow Analysis and Die Design 

Simulation of castings before die building has 

become a common practice in the last decade as a 

result of the increasing speed of computers and the 

accuracy predicted by mold flow programs such as 

High Strength Weight Reduction Materials 

MgO

AlSi

Figure 7. Individual gray-scale AES maps of Mg, 

oxygen, aluminum, and silicon. These maps correspond to the SEM in Figure 6. 

Table 3. Summary of secondary phases identified 

with X-ray diffraction with various performs and 

processes. PIVAC samples are in T4 condition. All 

others are as-cast 

Alloy/Process Secondary Phases Approx. MgO 

Identified w/ XRD particle size (nm) 

AZ91D/ CF25 

PIVAC MgO 11 +/- 2 

AZ91D/ CF10 

SiC15 Mg2Si MgO 10 +/- 2 

PIVAC 

AZ91D/ SiC25 CF3 

Squeeze Cast Mg2Si MgO 19 +/- 5 

AZ91D/ SiC25 CF3 

PLI Mg2Si MgO 79 +/- 5 

MAGMAsoft. MAGMAsoft casting simulation 

software, developed by MAGMA GmbH, of 

Aachen, Germany, is a widely used software package that enables foundries and casting users to simulate filling and solidification processes. The program 

allows for prediction of potential casting defects 

associated with incomplete filling and/or lack of 

feeding during solidification stages. The software is 

driven by an advanced set of algorithms based on 

laws of physics, thermodynamics, and computational fluid dynamics and is capable of simulating a 

wide variety of casting processes such as vertical/horizontal sand casting of ferrous and non18 
High Strength Weight Reduction Materials 

ferrous alloys; low-pressure, permanent mold and 

die casting of non-ferrous alloys; investment casting 

of ferrous and non-ferrous alloy, etc. The welldeveloped filling and solidification algorithms consider a variety of factors and parameters that may be 

used in a particular casting process. One such advanced feature of MAGMAsoft is the capability it 

provides to simulate filling of a runner system that 

contains a filter. 

The purpose of this activity is to simulate, using 

MAGMAsoft, the infiltration of a carbon fiber/SiC 

particulate preform with molten Mg. Using a modified squeeze cast process, the intent is to push the 

molten metal through the preform to produce a 

lightweight component with a high specific strength. 

Because of their high density, preforms obstruct the 

flow of molten metal in the mold cavity, often causing the formation of unfilled pockets and microporosity within the casting. Understanding flow 

mechanisms during infiltration is essential to obtaining optimum pouring temperature, filling rate, backpressure, and other parameters that define a robust 

process capable of consistently producing 100% infiltrated, sound Mg squeeze cast composites. 

The objectives of this portion of the work are as 

follows. 

Step 1. Explore the possibilities of using casting 

simulation software to model infiltration of preforms 

with Mg during the squeeze casting process. 

Step 2. Understand flow mechanisms during filling of the mold and attempt to predict potential defects that may result from incomplete infiltration of 

the preform. 

Step 3. Model different mold cavity configurations containing preforms of varying geometries. 

Results and conclusions obtained from the work 

described are communicated to GS Engineering and 

REL Machine to assist them with setup parameters 

during their experimental work. 

Step 1 

A typical setup of a squeeze cast process consists of a sprue through which molten metal is delivered either directly into the cavity or through runners, a mold cavity, and vents. In the Mg squeeze 

cast process explored by GS Engineering and REL, 

the sprue is positioned vertically (Z direction) and 

metal is delivered into the mold cavity from the bottom of the mold. The schematic of the setup is 

shown in Figure 8. 

FY 2006 Progress Report 

Figure 8. Typical setup of squeeze cast process. 

Based on the functionality description of 

MAGMAsoft, the program is fully capable of simulating a vertical casting process, including modeling 

the flow of metal through the filter. Therefore, it was 

thought possible to model infiltration of a preform 

during the filling stage of a squeeze casting process. 

In order to model the filling stage of the process, 

a certain amount of freedom was used with the 

MAGMAsoft package, including these examples: 

1. The formulas used in MAGMAsoft that describe 

the flow of metal through the filter are based on 

the pressure drop data vs stream velocity. The 

drop in pressure in the stream of metal as it 

flows through the filter largely depends of the 

permeability of the filter and its size. To run 

simulations, researchers at GS Engineering and 

REL experimentally derived pressure drop 

curves for each type of preform based on its 

density and size. These pressure drop curves 

were imported into the MAGMAsoft filter database to be used during experimental simulation 

runs. 

2. Modeling of the flow of metal through the preform-filter does not consider the complex internal topography of pores. In addition, the simulation is carried out assuming the absence of interaction between molten metal and the residual 

binder compounds remaining on the surfaces of 

preform fibers and particles. The potential lack 

of wetting of particles and fibers that may exist 

during the actual infiltration process is not ac19 
FY 2006 Progress Report 

counted for in the simulation because of the extreme complexity of mechanisms describing this 

phenomenon. 

Step 2 

The concept of modeling the infiltration of the 

preform in a squeeze casting process was tested in 

the following setup (Figure 9): 

1. A cylindrical cavity was set in a steel mold. 

2. A rectangular preform with a square crosssection was placed in the cavity. 

3. An inlet was placed at the bottom of the mold in 

immediate contact with the preform. 

4. A vent was added to the system to provide an 

outlet for gases pushed out of the mold cavity by 

the metal during infiltration. 

Figure 9. Infiltration of a rectangular 35% SiC–3% 

carbon fiber preform. 

Additional setup parameters used in the simulation are as follows: 

1. Mg alloy used in simulation: AZ91 

2. Alloy pouring temperature: 725°C 

3. Preform used: 35% SiC, 3% carbon fiber 

4. Temperature to which mold was preheated: 

300°C. 

5. Total fill time predicted by this simulation run: 4 

s. 

The results of the simulation showed several 

prevailing filling patterns: 

1. Because of the difference in cross-section geometry of the mold cavity and the preform, the 

setup allowed the molten metal to flow through 

the gaps between the mold and the preform. As a 

High Strength Weight Reduction Materials 

result, the alloy quickly enveloped the preform 

and entered the vent before complete infiltration 

of the preform. 

2. The infiltration of the preform proceeds from the 

bottom up. However, once the entire preform is 

enveloped by the alloy, additional infiltration is 

observed with the metal moving from the sidewall toward the centerline of the preform. 

The final stages of filling indicate that there exists a high potential for formation of center line porosity or areas of incomplete infiltration, as shown in 

Figure 10. 

Figure 10. Melt velocity distribution at 96% of fill 

time. 

Melt pressure distribution in the mold at 96% of 

filling supports this prediction—low melt pressures 

(14.9 ~ 76.1 psi) in the vicinity of a porosity pocket 

indicate that the flow of metal is reduced to the point 

of a complete halt. 

The main conclusion reached based on the results of this simulation is that the gap between the 

mold and the preform allows a fraction of the melt 

front to follow the path of least resistance. This creates a condition in which the metal envelopes the 

preform long before it is fully infiltrated. Subsequent infiltration of the mold may be impeded by the 

gases trapped inside the preform. Following this 

modeling run, it was recommended that a setup be 

explored in which the gap between the mold and the 

preform is eliminated by matching the cavity of the 

mold to the geometry of the preform, i.e the final 

geometry of the casting. 

20 
High Strength Weight Reduction Materials 

Step 3 

The setup modeled in Step 3 is depicted in Figure 11. The mold cavity precisely matches the geometry of a preform that when fully infiltrated 

would result in a Mg squeeze-cast connecting rod 

casting. The goal of the setup is to force the metal 

directly through the preform by eliminating any gaps 

between the mold and the preform. This simulation 

was carried out using the same process parameters as 

those in step 2. The stages of the filling sequence are 

shown in Figure 12. 

Figure 11. Mold set up to squeeze-cast a preform 

reinforced AZ91 Mg connecting rod casting. 

Filling of casting reinforced by preform 

The results of this modeling run demonstrate a 

generally good pattern of filling. The entire crosssection of the preform was filled without the formation of visible pockets of porosity. The problem 

arises at the final stages, around 94 to 100% of filled 

volume. A large pocket of porosity was formed at 

the very top of the connecting rod casting, as shown 

in Figure 12. 

Full infiltration is observed throughout the volume of the preform up to ~94% of fill. However, 

beyond this point, an apparent pocket of porosity is 

formed at the top of the rod casting. Higher melt 

pressures noted around the pocket point to the fact 

FY 2006 Progress Report 

Figure 12. Mold flow simulation 94% complete 

showing potential filling issues. 

that the flow of metal is slowed or halted by trapped 

gases. 

The distribution of melt pressure results provides another valuable bit of information on the potential problems that may arise during the infiltration 

stage. A rapid increase in back-pressure is observed 

in the sprue as infiltration proceeds beyond 50% of 

total fill. Pressure in the sprue reaches as high as 

23,730 psi at 96% of total fill. It is suggested that 

these pressure values be compared with the compression strength of the preform to make sure that 

the preforms are not cracked or crushed during infiltration. The total fill time of this system was predicted at 2.65 s. 

Based on the velocity and pressure results, it is 

believed that the lack of infiltration at the top of the 

casting is caused by incorrect design of the overflow/venting system. This flaw can be fixed by a 

relatively simple redesign of the venting system. It is 

suggested that two of the three overflow outlets be 

eliminated, leaving only one outlet at the top of the 

casting. This outlet would connect with a large overflow reservoir that would branch into two vertical 

vents. 

Also, the contact with the sprue at the bottom of 

the casting was widened to increase the amount of 

metal entering the mold. The modified mold layout 

is shown in Figure 13. 

The setup with the suggested modification in the 

venting system was simulated to verify that the 

change would improve infiltration at the top of the 

casting. The casting process parameters were kept 

unchanged to maintain consistency between simulations. Infiltration progressed through the cross21 
FY 2006 Progress Report 

Figure 13. Modified layout to improve venting. 

section unobstructed and completed at the top of the 

connecting rod casting without any indications of 

potential porosity pockets. 

Similarly, pressure distribution results did not 

contain any signs that would mark potential insufficient infiltration. However, the back-pressure in the 

sprue rapidly rose as the melt front moved through 

the preform to the top of the casting. Steps may be 

needed to measure the actual back-pressure during 

experimental runs, and the head pressure may need 

to be adjusted to prevent damage to the preform. 

Computer modeling of fiber/particulate pre-form 

infiltration with molten Mg using MAGMAsoft provided sensible results that gave insight into the filling mechanisms taking place during the filling stage 

of the process. The following conclusions were 

drawn from the simulation results: 

1. Infiltration may be improved when the preform 

comes into direct contact with the mold wall, 

thus eliminating gaps and forcing the metal to 

flow through the preform instead of around it. 

2. The back-pressure of the melt built up in the 

sprue may need to be controlled to avoid exerting excessive compressive forces on the preform 

that might damage its structural integrity. 

3. The venting system is a critical element to successful infiltration. It must be designed to avoid 

High Strength Weight Reduction Materials 

entrainment of gases that will impede the progress of metal through the preform. 

It is highly recommended that each setup modeled with MAGMAsoft be experimentally tested to 

draw correlations between results. 

Die Fabrication and Casting Trials 

From the iterations in the MAGMAsoft model, a 

good feel for the metal flow through the preform 

was realized. These steps eliminated costly die reworking, as the design of the die was altered to enhance metal flow. The die, shown in Figure 14, is 

the result of a collaborative effort ranging from the 

die designer to the mold flow analyst to the machinist cutting the final dies. 

ProEngineer Model MAGMA moldflow 

Simulation 

ConRod Die, 

Fabricated 

Die Design 

Moldflow 

Analysis w/ 

preform 

Die and preform fabrication 

Figure 14. The process diagram of the connecting rod 

die build. 

Aluminum connecting rods were the first components cast in the new die to ensure a good seal 

was obtained before switching to the more volatile 

metal Mg. Initial experiments were performed. The 

final casting is seen in Figure 15. 

Validation of Model with Experiments 

Model validation is ongoing, but some preliminary data indicate that the model predictions and 

casting observations do agree with each other. CT 

scans of a conrod cast with an aluminum alloy indicate good filling of the cavity (Figure 16). Issues 

arise at the thick-to-thin transitions, as is expected in 

many castings. These issues can be overcome with 

adjustment of casting process conditions. 

Press Validation 

A series of infiltration trials were cast on the 

GSE press to narrow process parameter windows 

22 
High Strength Weight Reduction Materials 

Figure 15. Initial magnesium 

connecting rod casting in die. 

A-A B-B C-C D-D E-E F-F 

A-A B-B C-C D-D E-E F-F 

Figure 16. CT scan of a reinforced connecting rod 

showing the degree of aluminum infiltration of the 

preform. 

0 

5 

10 

15 

20 

25 

30 

35 

A356-T7 A356/SiC28-T7 Material 

Modulus/UTS 

Figure 17. Comparison of A356 and A356-33SiC. 

and to compare mechanical properties achieved with 

listed data on MMCs. Figure 17 indicates strength 

and stiffness values achieved in the casting of A356 

into a 33 vol % SiC preform. It also compares the 

A356/SiC33-T7

Modulus (Msi) 

UTS (ksi) 

FY 2006 Progress Report 

pure alloy with the MMC. A 36% increase in stiffness and an 18% increase in ultimate tensile strength 

was noticed with the reinforcement. 

Initial casting with A356 and A356-33SiC 

showed the mechanical properties achieved compare 

very well with published data. That raises our confidence in our casting system. Our initial trials with 

Mg look promising, as we were able to infiltrate our 

CF25 preforms with some success. This work will 

continue in the coming months, as our molten metal 

transfer system is currently functional. 

Conclusions 

Process control in Mg MMCs is complicated by 

the reactive nature of the molten metal. Interactions 

between the metal and the preforms can be positive 

for good fiber-matrix bonding, but too much of these 

reaction products can produce unwanted intermetallics and oxides in the matrix. The reaction products 

of the Mg alloy reacting with the silica binder have 

been tabulated and the amounts can be seen from the 

X-ray diffraction data. The quantities of Mg2Si and 

MgO change with the casting speed and solidification times; therefore, contact time between the molten metal and the preform must be optimized to obtain good bonding between the two without degradation of the reinforcement or the overproduction of 

excess constituents. In light of these observations, 

process conditions are precisely controlled to optimize the casting of selectively reinforced Mg composites and will be reported in future publications. 

Future Work 

Future work will focus on refining the materials, 

i.e., matrix/reinforcement, and process parameters 

for a production-capable casting system through a 

series of casting trials on the GSE casting press with 

microstructural analysis of micro and macro interfaces. Mechanical testing will be performed in bench 

and field testing, which will give feedback to optimize process and materials based on those results. 

The Gantt chart in Figure 18 details the work and 

timeline for year 2 of the project. 

Acknowledgements 

The authors would like to thank DOE and Oak 

Ridge National Laboratory (ORNL) for the funding 

23 
FY 2006 Progress Report 

Figure 18. Project plan for remaining work to be 

completed. 

to perform this research. Also, thanks are due to 

Harry Meyer of ORNL for the AES analysis of cast 

samples, Edward Laitila (Michigan Technological 

University) for assistance in performing the X-ray 

diffraction analysis, and GOM Technologies for the 

pressureless infiltration trials. 

High Strength Weight Reduction Materials 

Presentations/Publications 

R. Hathaway, A. Loukus, C. Johnson, T. Wood, 

J. Loukus, A. Halonen, G. Simula, V. Pikhovich, B. 

Coleman, and D. Weiss, “Manufacturing Process 

Influence on Microstructural Features of Selectively 

Reinforces Magnesium Metal Matrix Composites,” 

in TMS Proceedings, San Antonio, 2006. 

J. E. Loukus, A. Loukus, A. Halonen, “The Effects of Processing Methods on the Mechanical 

Properties of Cast Magnesium Metal Matrix Composites,” AFS International Conference on High Integrity Light Metal Castings, Indianapolis, October 

31–November 1, 2005. 

24 
High Strength Weight Reduction Materials FY 2006 Progress Report 

3. MATERIALS PROCESSING TECHNOLOGIES 

A. Development of an Advanced Squeeze Casting Process for Producing HighIntegrity Truck Components 

Principal Investigator: David C. Weiss 

Eck Industries, Incorporated 

1602 North 8th Street, Manitowoc, WI 54221-0967 

(920) 682-4618; fax: (920) 682-9298; e-mail: dweiss@eckindustries.com 

Chief Scientist: James J. Eberhardt 

(202) 586-9837; fax: (202) 587-2476; e-mail: james.eberhardt@ee.doe.gov 

Field Technical Manager: Philip S. Sklad 

(865) 574-5069; fax: (865) 574-6098; e-mail:skladps@ornl.gov 

Contractor: Eck Industries, Incorporated 

Contract No.: 4000022893 

Objective 

• Develop the equipment and process technology for an advanced squeeze casting (ASC) process to enable production of high-integrity cast metal components. 

— Integrate the advantages of two casting methods—low-pressure permanent mold (LPPM) and direct 

squeeze casting —to attain non-turbulent fill of the die, high solidification rates to refine microstructural 

features, and solidification under pressure to minimize microporosity. 

— Design and build a new kind of casting machine and develop process technologies needed to cast highintegrity truck components from nonferrous alloys. 

Approach 

• Design and construct a casting machine that integrates a low-pressure metal delivery system suitable for either 

aluminum or magnesium alloys, reliable gate shut-off technology, and direct application of squeeze pressures 

up to 103 MPa (15,000 psi). This machine is intended for production. 

• Develop a gate shut-off mechanism that will operate reliably in a production environment. 

• Select a casting that will be produced in the casting machine built for this project. 

• Use fluid-flow and solidification analysis to predict optimum flow conditions for metal entering the mold and 

differential squeeze pressure requirements. 

• Design and build cast tooling 

• Develop process technologies for the low-pressure/squeeze casting of aluminum alloys and evaluate the effect 

of various process parameters on casting integrity. 

• Apply developed technologies to the production of a selected automotive component. 

25 
FY 2006 Progress Report High Strength Weight Reduction Materials 

Accomplishments 

• Designed, built, and installed ASC machine at Eck Industries. 

• Designed and fabricated squeeze cast tooling . 

• Developed process parameters for prototype production of B356-T6 connecting rods. 

• Produced ASC and LPPM castings for determination of static and dynamic properties of ASC-processed B356-

T6. 

Future Direction 

• Determine static and dynamic properties of ASC-processed B356-T6. 

• Demonstrate production viability of ASC casting machine and process. 

Introduction 

Squeeze casting is the solidification of liquid 

metal under pressure in a closed metal die. The casting produced has improved properties and a more 

uniform microstructure compared with those produced by traditional molten metal fabrication techniques. The improved properties are achieved 

through controlled entry of the metal into the die 

through large gates, which reduces turbulence, and 

high solidification rates with resulting refinement of 

microstructural features. 

There are two squeeze casting methods—direct 

and indirect. The method and efficiency of pressure 

application distinguish the two processes. Direct 

squeeze casting applies pressure directly on the casting. Indirect squeeze casting applies pressure indirectly via the gating system. The length of the gating 

system and any partial solidification prior to complete solidification of the part reduces the magnitude 

of the pressure applied and may result in casting porosity. For the indirect method, the dimensions of 

the die cavity control part dimensions. For the direct 

method, the part dimensions are controlled by the 

die cavity dimensions and the finish position of the 

top die. Thus direct squeeze casting requires precise 

machine control to attain the desired part dimensions. In addition, the direct method requires in-gate 

shut-off technology to control pressure during cavity 

pressurization. Both methods suffer from poor metal 

handling from the furnace into the mold or gating 

system. Inadequate control of pouring speed generates turbulence and casting defects such as oxide 

inclusions that reduce fatigue properties. 

Earlier attempts to improve the squeeze casting 

process by combining the advantages of lowpressure die fill with those of squeeze casting demonstrated substantial mechanical property improvement and reduced component porosity. However, 

continuous production operation of the equipment 

was not achieved. The machines used were not specifically designed to meet the needs of the process. 

Technical Approach 

The results of this project will provide (1) a production-viable ASC machine, (2) process technology 

that will improve the strength and reliability of cast 

automotive components, and (3) demonstrated production viability of the equipment and process. 

Machine design. Previous efforts demonstrated 

that modification of current cast machine designs 

would not provide a production-capable ASC machine. An assessment of the capabilities and limitations of current machine designs demonstrated that 

• Low-pressure cast equipment typically does not 

have the structural rigidity needed to withstand 

the high pressures required of squeeze casting. 

• Die-casting and currently available squeeze casting machines are not designed for proper (nonturbulent) molten metal entry into the die cavity. 

• Neither type of equipment is typically available 

with the hydraulic and electronic controls necessary to attain the die cavity motions required for 

the direct squeeze casting process. 

After much consideration, the project team decided it was easier and more cost-effective to build a 

machine from the ground up than to modify an existing cast machine design. To facilitate the design and 

construction, the team partnered with Empire Castings, Inc., an experienced producer of low-pressure 

casting equipment. 

26 
High Strength Weight Reduction Materials 

Process development. Bendix, Inc., is an industrial manufacturer of truck components with an interest in reducing manufacturing cost by the use of 

high-quality castings. Consultation with Bendix led 

to the selection of an air compressor connecting rod 

for prototype production in the ASC machine. This 

component has a level of complexity that can demonstrate the production capability of the ASC machine and process, show the need for mechanical 

properties higher than those attainable with commercially available castings, and provide an opportunity for production with the ASC process. 

The cast part will be modeled for fluid flow and 

solidification. This work will be done in parallel 

with equipment design. Simulation will provide an 

understanding of optimum metal flow conditions 

and squeeze requirements, if any, to contend with 

solidification shrinkage. This is very important to 

the success of the program and, potentially, to the 

details of the machine design. The ASC machine 

will have two mechanisms to feed the casting—lowpressure, nonturbulent fill of the die cavity and highpressure, direct squeeze to minimize solidification 

shrinkage. Modeling experiments will explore the 

interaction of the two feed mechanisms and select 

process parameters that lead to process optimization. 

Modeling results will be used to refine the die design and select initial casting process parameters. 

After tooling construction and ASC machine delivery, castings will be made under the range of 

process conditions anticipated during modeling. The 

alloy to be used for development is the cast alloy 

B356. Evaluation of those castings will include radiography, die penetrant inspection, and tensile testing. It is anticipated that this part of the program will 

require multiple iterations to produce the desired 

product quality. Some equipment and tooling modification may be required during this part of the program. 

After successful production of development 

castings, the production viability of the equipment 

and process will be demonstrated by 

• Production of a connecting rod with reduced 

porosity and improved mechanical properties 

compared with the current die-cast connecting 

rod used by Bendix. 

• Continuous production with the ASC machine 

and connecting rod tooling to include at least 

one 5-h continuous run with the equipment running at least 3 days in a week. 

FY 2006 Progress Report 

The goal of the project team is to implement 

commercial part production on an ongoing basis. 

Program Schedule 

The revised scheduled start and end dates for the 

various program tasks are listed in Table 2. 

Table 2. Overall program schedule (revised) 

Program Task Start date Finish date 

Machine design June 2003 Feb. 2004 

Machine build Mar. 2004 Nov. 2004 

Tooling design July 2003 Oct. 2003 

Tooling build Nov. 2003 Mar. 2004 

Process model Oct. 2004 Dec. 2004 

In-gate SO development Nov. 2003 Dec. 2003 

Machine install Nov. 2004 Nov. 2004 

Casting development Dec. 2004 Dec. 2005 

Evaluation and testing July 2006 Dec. 2006 

Final report Nov. 2006 Jan. 2007 

Program Status 

ASC machine. The 600-ton-capacity ASC casting machine is installed at Eck Industries (Figure 1) 

and has been used to cast tensile bar specimens and 

prototype connecting rods (Figure 2). 

Data acquisition system. The computerized 

data acquisition system has worked flawlessly. The 

pressure, displacement, and temperature data have 

been used to understand the operation of the ASC 

process and establish process parameters. A plot of a 

typical data set is shown in Figure 3. 

Process development. Modification of the machine’s PLC program to use timers instead of temperature data for controlling start and stop of 

machine operations was proved successful by the 

sequential accomplishment of 40 castings. A typical 

27 
FY 2006 Progress Report High Strength Weight Reduction Materials 

Figure 1. ASC machine installed at Eck 

Industries. 

Figure 2. Typical casting from 40 shot run. 

Figure 3. Typical data obtained from data 

collection system during casting runs. 

casting is shown in Figure 2. Subsequently the 

four-cavity mold designed and fabricated for the 

program was used to make two sets of castings—one 

set by squeeze casting and the other via the conventional permanent mold process. The two sets of tensile bar castings were machined and are being used 

to determine tensile and fatigue properties of B356-

T6 castings. Subsequent data analysis, fractography 

and image analysis will determine the quality and 

property enhancement attainable to the ASC process. 

Conclusions 

Consistent operation of the ASC machine has 

been obtained. Fatigue and tensile specimens have 

been machined and evaluated for quality level via 

radiography. The determination of static and fatigue 

properties at room temperature has been started and 

completion is scheduled for December 2006. As 

shown in Table 2, the schedule has been revised to 

reflect expected completion dates. 

Prototype Production for Industry 

Eck Industries is working with an industrial customer to determine the viability of the ASC process 

for the production of a transmission hub (see Figure 4). When it is produced via die casting, the rejection rate for this part is higher than desired. 

Figure 4. Transmission hub for automotive 

 application. 

Casting simulation predicts that both die and 

LPPM casting will contain porosity in areas that 

need to be “porosity free.” Use of the ASC process 

is expected to provide the desired “porosity free” 

casting, reduce the rejection rate, and reduce overall 

manufacturing costs. Lower porosity levels are also 

28 
High Strength Weight Reduction Materials 

expected to increase the fatigue life of the component. Prototype casting of this component is scheduled for early December 2006. Successful prototype 

casting is expected to lead to a production contract 

for this and other similar components. 

FY 2006 Progress Report 

Presentations and Publications 

D. C. Weiss, “Design and Fabrication of a Machine for an Advanced Squeeze Casting (ASC) 

Process,” presented at the AFS International Conference on High Integrity Light Metal Castings, Indianapolis, October 31–November 1, 2005, and 

published in Proceedings of the AFS International 

Conference on High Integrity Light Metal Castings, 

American Foundry Society, Schaumberg, IL, 2005. 

29 
FY 2006 Progress Report High Strength Weight Reduction Materials 

30 
High Strength Weight Reduction Materials FY 2006 Progress Report 

B. Counter-Gravity (Hitchiner) and Pressure-Assisted Lost Foam Magnesium 

Casting 

Principal Investigator: Qingyou Han 

Oak Ridge National Laboratory 

Oak Ridge, TN 37831-6083 

(865) 574-4352; fax: (865) 574-4357; e-mail: hanq@ornl.gov 

Chief Scientist: James J. Eberhardt 

(202) 586-9837; fax: (202) 587-2476; e-mail: james.eberhardt@ee.doe.gov 

Field Technical Manager: Philip S. Sklad 

(865) 574-5069; fax: (865) 574-6098; e-mail: skladps@ornl.gov 

Participants: 

Ralph B. Dinwiddie, Oak Ridge National Laboratory 

Pete Angelini, Oak Ridge National Laboratory 

Graham V. Walford, Walford Technologies 

Ken Currie, Tennessee Technological University 

Mohamed Abdelrahman, Tennessee Technological University 

Fred Vondra, Tennessee Technological University 

Contractor: Oak Ridge National Laboratory 

Contract No.: DE-AC05-00OR22725 

Objective 

Evaluate the Hitchiner process and pressure-assisted lost foam casting of magnesium (Mg) alloys. In particular, 

• Develop the casting system. 

• Develop low-density foams and appropriate coatings. 

• Establish casting parameters for selected Mg alloys. 

• Conduct computer simulations of the process to be used to generate design guidelines. 

• Validate the process by characterizing microstructure and properties of castings. 

Approach 

• Compare lost foam casting characteristics of Mg and aluminum (Al) at 1 atmosphere 

• Castability/integrity/mechanical properties 

• Compare lost foam Mg casting with Hitchiner lost foam Mg casting and pressure-assisted casting 

— Castability/integrity/mechanical properties 

• Document the features of the Hitchiner lost foam Mg casting 

— Possible applications, typical features of casting that are suitable for the Hitchiner process, and limitations 

of the Hitchiner process 

31 
FY 2006 Progress Report High Strength Weight Reduction Materials 

Accomplishments 

• Developed melting and casting capabilities at Tennessee Technological University (TTU) and Oak Ridge National Laboratory (ORNL). 

• Observed significant difference between Mg lost foam casting and Al lost foam casting in mold filling and molten metal/foam interaction. 

• Developed a system for the selection of foam materials and coatings. 

• Carried out experiments to develop baseline data and to evaluate the Hitchiner process for lost foam casting 

Future Direction 

• Select/develop coatings and foams for the Hitchiner process for lost foam magnesium casting. 

• Complete the design of a parametric study of the parameters affecting the castability of Mg alloys under the 

Hitchiner process conditions. 

• Start the simulation of the Hitchiner/lost foam process. 

• Document the features of Hitchiner lost foam Mg casting. 

Introduction 

The last 10 years have seen dramatic increase 

in the production and utilization of Mg. This demand is primarily due to an increase of 20% annually in the use of Mg in the automobile industry 

driven by the desire to increase vehicle fuel efficiency by reducing vehicle weight. Magnesium 

has found uses in several automotive components 

because of its low density and high strength-toweight ratio compared with conventional automotive materials. More recently, Mg castings have 

been introduced into pickup trucks and vans, such 

as the standard size Ford F-150. Additional opportunities exist for the use of Mg casting in these 

vehicles, as well as in heavier trucks. Even with 

these successes, however, Mg has not been as 

widely accepted as it might have been for the following reasons: the cost of the alloys, the lack of 

long-term field validation or controlled fleet testing data for powertrain components, and lack of a 

technological base for foundry casting of Mg alloys. 

Magnesium castings are traditionally produced 

by die casting or precision sand casting. Each of 

these processes has its advantages and disadvantages for producing components. The major disadvantage of conventional die casting processes is 

that certain geometries cannot be produced using 

steel die sections and are limited by moving cores. 

In addition, die casting often results in unacceptable levels of porosity. On the other hand, precision sand casting can produce a wide range of 

shapes with high levels of soundness, but at an 

unacceptable cost. Although successful castings 

can be manufactured, continuing problems experienced during the casting of the Mg alloys include 

porosity formation, oxidation, and hot tearing/cracking formation. 

In order to take greater advantage of the attractive properties of Mg, alternative approaches 

that promise to be more robust and lead to the development and fabrication of larger, thinner wall 

castings must be considered. Lost foam castings 

can be competitive with other casting techniques 

for applications with complex internal passages 

(coring) and can reduce machining, save space, 

and/or integrate multiple complex features or 

components into a single casting. 

Previous studies at ORNL on gravity lost 

foam casting of Al have determined that many 

factors affect the quality of the casting outcome. In 

particular, the metal filling process of the lost 

foam pattern is greatly influenced by the pattern 

properties. The advancing metal front decomposes 

the foam into gaseous and liquid decomposition 

products that must be eliminated from the casting 

volume to complete the process. Pattern density 

and the pattern gas permeability arising from the 

combination of the foam and ceramic coating are 

critical factors in the process, yet they are not 

32 
High Strength Weight Reduction Materials 

measured on any commercial process line. Variances in these properties are difficult to control 

and result in variance in the metal filling behavior. 

The physical properties of Mg would increase this 

problem further. In particular, the low density, 

latent heat, and high reactivity of the molten Mg 

alloy are serious issues for its castability under lost 

foam casting conditions. 

These problems can be overcome by making 

use of pressure-/vacuum-assisted mold filling together with closed-loop process control. The 

Hitchiner process, which uses vacuum as a driving 

force for mold filling, is a logical choice for lost 

foam casting of Mg alloys. The change of pressure 

across the advancing metal front can be used to 

accommodate variances in the gas permeability of 

the foam and coating for the whole pattern and for 

local variations in the pattern. Pressure can also be 

used to assist in mold filling. 

The challenges of using the Hitchiner process 

and pressure-assisted lost foam Mg casting are in 

the areas of setting up the experimental equipment 

for the research; parametric study of the parameters affecting the processes; development of lowdensity foams for the casting of Mg alloys; control 

of permeability of the sand, foam and foam coatings; and the use of controlled atmospheres for 

reducing metals reactance, chemistry, and the 

thermal profile of the pattern to be cast. This article is the second-year report for this project. The 

goal of the second-year research was to develop 

the Hitchiner casting process and evaluate this 

process for Mg casting. 

The Casting Systems 

The existing metalcasting facilities at TTU 

have been modified for handling molten Mg alloys 

and for using the Hitchiner process for lost foam 

casting of Mg alloys. A Thermtronix solid state 

Al/ Mg electric resistance furnace has been purchased and facilitated for use. It has a 200–300 lb 

charge capability (depending on whether it is used 

for Mg or Al), and a special-purpose cover made 

from mild steel has been added to the furnace to 

allow for introduction of a cover gas to prevent the 

molten Mg alloys from reacting with the atmosphere. The cover also has a hinged lid on the very 

top to facilitate direct insertion of the countergravity sprue. The Hitchiner counter-gravity casting system, shown in Figure 1, has been successFY 2006 Progress Report 

fully tested for Mg castings and Al castings. Both 

a simple flat plate pattern (provided by GM) and a 

widget pattern (provided by Foseco Morval) were 

successfully cast using Mg alloy AM60B and Al 

alloy 319. A 1% sulfur hexafluoride/99% carbon 

dioxide cover gas was used over the molten Mg 

either in the crucible or in the furnace. A set of 

safe operating procedures were adopted to safely 

handle and cast Mg alloys in the current casting 

facility. 

Figure 1. The Hitchiner casting system is being tested 

by TTU and ORNL researchers for lost foam casting of 

Mg alloys. 

Foam Development and Selection 

A broader study of foam decomposition has 

been undertaken that seeks to develop a kinetic 

and thermochemical model that describes polystyrene foam decomposition for a broad range of 

heating rates for oxidizing and non-oxidizing environments. Understanding the foam decomposition 

processes under different process conditions, including foam decomposition in the presence of 

molten Mg, served as the motivation behind this 

work. This effort might help in improving casting 

quality in ways such as reducing residue entrainment. 

Based on a thorough literature search on appropriate kinetic models applicable to the polystyrene foam decomposition phenomena encountered 

in the lost foam casting process, experiments have 

been designed and carried out to investigate the 

foam degradation kinetics under the process conditions. The main purpose of this experimentation 

is to obtain data on the composition of polystyrene 

33 
FY 2006 Progress Report 

foam degradation products as a function of melt 

temperatures and processing times. 

Initially, a simple burn chamber using an infra-red heat source was constructed for these experiments, but a more precise reactor concept was 

conceived. Figure 2 shows the burn chamber used. 

The reactor consists of a heater maintained at a 

constant surface temperature and an automatic 

feeder that feeds the foam at a velocity “v”’ from 

the automatic feeder, which simulate the actual 

metal temperature and melt flow rate, respectively, 

in the commercial casting process. When the foam 

impinges on the heater surface, it decomposes and 

forms various products. These liquid and gaseous 

products are collected and analyzed for their composition. Decomposition data could be collected as 

a function of both foam push rate and metal temperature. A schematic of the proposed reactor is 

shown in the figure. 

Figure 2. The burn chamber for characterizing foam 

decomposition. 

In addition to analyzing and modeling the 

foam decomposition products, another important 

aspect is the characterization of the foam pattern. 

Walford Technologies, Inc. (WTI) has expanded 

its pattern quantification system to provide measurement of density, sonic propagation surface finish, and dimension profiles for patterns up to 36 

in. in length. Figure 3 shows the setup of the infrared camera at a molding operation. The 

High Strength Weight Reduction Materials 

Figure 3. Infrared camera installation at a molding 

operation. 

camera is targeted to observe the expanded polystyrene (EPS) foam part. 

Further work is required to correlate the properties of the EPS part as measured at the time of 

molding with the properties of a part that is suitable for casting. However, effective control of any 

process, and in particular of the process of using 

EPS in Mg casting operations, requires real time 

control and assessment of the part. This has not 

been previously accomplished. 

Experimental Results on Castability 

Castability experiments have been carried out 

at ORNL and TTU. Three types of patterns were 

used: a simple flat plate, the GM window pattern, 

and a widget pattern. The plate pattern was used 

for observing mold filling and for evaluating porosity distribution. The window pattern and the 

widget pattern were used for evaluating the castability of alloys. Five types of foam materials from 

Styrochem were tested: type A (T-17-170, ventless 

tool), type B1 (T-185, vented tool), type B2 (T185, ventless tool), type C (T-170, vented tool), 

and type E (T-185 with bromated additive, ventless tool). A high-permeability coating was used 

for the testing. Castings of an A356 alloy and an 

AM60B alloy were made. 

Casting temperature has a significant effect on 

the castability of the Mg AM60B alloy. Castings 

were made using the GM window patterns made 

of B1 foam. A cold shut defect occurred at the 

corner where the metal filled last when the pouring temperatures were around 1400 to 1450°F. 

Increasing the pouring temperature reduced the 

34 
High Strength Weight Reduction Materials 

cold shut defect. When the pouring temperature 

was 1500°F, the cold shut defect was eliminated. 

However, 1500 °F is about 355°F higher than the 

liquidus temperature of the alloy. At such a high 

temperature, oxidation and dross formation can be 

a significant problem. A similar casting can be 

made at 1300°F using the Al alloy A356, indicating it is more difficult to cast Mg alloys than Al 

alloys. 

The Hitchiner process enhances the castability 

of Mg alloys significantly. Figure 4 shows a widget Mg casting made using the Hitchiner process. 

The casting was cast at 1300°F and had no cold 

shut defect. Considering that the widget casting is 

more complex in geometry than a window casting, 

a higher pouring temperature would be required in 

order to make a casting without cold shut defect if 

the widget casting were made under the normal 

lost foam casting conditions. The use of the 

Hitchiner process makes it possible to produce 

geometrically perfect castings at a much lower 

pouring temperature. The results also suggest opportunities for using this process for making a 

casting larger than that made by using the normal 

lost foam casting process. 

Figure 4. A geometrically perfect Mg alloy lost foam 

casting made using the Hitchiner process. 

A systematic research plan was developed using a mixed-level Taguchi design program and a 

random number generator. The research evaluates 

the effects of the key parameters or factors of the 

Hitchiner process on the castability and mechanical properties of AM60B alloys. These parameters 

include the pouring temperature, the pressure or 

FY 2006 Progress Report 

vacuum, the coating, and the types of foams. The 

experimental studies are still under way. 

Experimental Results for Porosity Formation 

The porosity distributions in flat plate castings 

of A356 and AM60B made using the lost foam 

casting process were characterized. The effect of 

degassing on the porosity distribution in the flat 

plate casting was also investigated. Castings of 

A356 alloy and AM60B alloy were made using a 

flat plate pattern of 76.2×152.4 mm and 12.7 mm 

thick (3×6 in. and 0.5 in. thick). The use of a simple geometry plate casting eliminated the effect of 

geometry on porosity formation. The casting was 

bottom-gated. A large fiber downsprue was used, 

which provided the pressure head and served as a 

riser for feeding the macro shrinkage. The flat 

plate castings were cut into small specimens. Each 

specimen was used for density measurement using 

the apparent density measurement method. 

The measured density distribution of the 

AM60B plate casting is shown in Figure 5a and 5b 

for degassed and non-degassed molten metal, respectively. The density distribution in the AM60B 

castings is fairly uniform, varying between 1.775 

and 1.778 g/cm3

. This small variation in the density data is within experimental error. Degassing 

did not affect the density or porosity formation 

much. No interdendritic pores were found in the 

casting, indicating the downsprue was large 

enough to avoid the formation of macro-porosity. 

Figure 6a and 6b illustrate the density distribution in the flat plate casting of A356 with and 

without rotary degassing. The in-gate location is in 

the middle at the right side of the diagram (x=35 

mm and y=0 mm). Comparing Figure 6a and 6b, 

the density of the degassed casting is higher than 

that of the un-degassed one. The density value 

varies from 2.634 to 2.654 g/cm3

 for the degassed 

casting, but it varies from 2.60 to 2.625 g/cm3

 for 

the un-degassed casting. The density data strongly 

suggest that degassing is an effective means of 

reducing porosity in Al lost foam castings. 

The trend of porosity distribution in both castings, degassed and un-degassed, is identical. The 

local density of the alloy was higher near the gate 

35 
FY 2006 Progress Report High Strength Weight Reduction Materials 

(a) (a) 

(b) 

Figure 5. The measured density distribution in Mg 

AM60B alloy castings (a) degassed, and (b) undegassed. The casting was gated at the middle on the 

right side of the plot. 

and lower away from the gate. The lowest-density 

region occurred in the middle near the end of the 

plate casting. A significant amount of porosity 

existed in this region. Figure 7 shows the microstructure of the Al plate casting near the region 

where the density is the lowest. The pores there 

are interdendritic, rather than isolated spherical as 

usually seen in the regions near the gate of the 

casting. 

The mechanism of the porosity formation in 

Al lost foam casting was investigated using an 

infrared camera to observe mold filling during the 

lost foam process. The metal advanced faster near 

the edge of the plate casting because the decomposition product of the foam could easily escape 

from the edges. As a result, the flow front was Vshaped. The low-porosity region occurred at the 

center of the V-shape partly because of entrapped 

(b) 

Figure 6. The measured density distribution in Al 

A356 castings (a) degassed, and (b) un-degassed. The 

casting was gated at the right side. A low-density region was identified near the left side of the Al casting. 

foam residues and partly because the metal front 

in that area collected more hydrogen because it 

was in contact with the decomposition atmosphere 

for a longer time than was the metal near the 

edges. 

This study on porosity distributions in Mg alloy and Al alloy lost foam casting showed that 

porosity formation is less an issue in Mg lost foam 

casting than in Al lost foam casting. Most likely 

that is due to the fact that hydrogen has a higher 

solubility in the solid Mg alloys than in the solid 

Al alloys. Less porosity formation is one of the 

benefits of using the lost foam process for making 

Mg castings. 

Conclusions 

The objectives of the second-year research in 

this project have been met. Melting and lost foam 

36 
High Strength Weight Reduction Materials 

Figure 7. Interdendritic porosity occurred in the region 

where the density is the lowest on the Al plate casting. 

casting capabilities have been developed at TTU 

and ORNL. The Hitchiner process has been established at TTU for testing counter-gravity and pressure-assisted lost foam Mg casting. A burn 

chamber has been modified for studying polystyrene foam decomposition for a broad range of 

heating rates for oxidizing and non-oxidizing environments. 

High-quality lost foam castings of Mg and Al 

alloys have been made using both the conventional lost foam and the Hitchiner processes. The 

castability of Mg alloys under Hitchiner conditions (vacuum-assisted lost foam conditions) is 

much better than that under conventional lost foam 

casting conditions. 

FY 2006 Progress Report 

The porosity distribution in lost foam casting 

of Mg alloy is significantly different from that in 

Al alloys. Uniform distribution of porosity is observed in lost foam Mg casting, whereas the porosity distribution in Al alloy casting is highly 

nonuniform. The application of these experimental 

results will be explored in the third-year research. 

Publications 

Q. Han and H. Xu, “Fluidity of Alloys under 

High Pressure Die Casting Conditions,” Scripta 

Materialia, 53, 7–10 (2005). 

Q. Han, R. Dinwiddie, P. K. Sklad, K. Currie, 

F. Vondra, M. Abdelrahman, G. Walford, D. 

Nolan, and T. Nedkova, “Lost Foam Casting of 

Magnesium Alloys,” in Magnesium Technology 

2007, TMS, in press. 

Q. Han, R. Dinwiddie, P. K. Sklad, K. Currie, 

F. Vondra, M. Abdelrahman, G. Walford, 

D. Nolan, and T. Nedkova, “Comparison of Lost 

Foam Casting of AM60B Alloy and A356 Alloy,” 

Transactions of AFS 2007, in press. 

37 
FY 2006 Progress Report High Strength Weight Reduction Materials 

38 
High Strength Weight Reduction Materials FY 2006 Progress Report 

C. Wrought Magnesium Alloy/Process Development 

Principal Investigator: J. A. Horton 

Oak Ridge National Laboratory 

P.O. Box 2008, Oak Ridge, TN 37831-6115 

(865) 574-5575; fax: (865) 574-825; e-mail: hortonja@ornl.gov 

Principal Investigator: S. R. Agnew 

University of Virginia 

116 Engineer’s Way 

P.O. Box 400745, Charlottesville, VA 22904-4745 

Chief Scientist: James J. Eberhardt 

(202) 586-9837; fax: (202) 587-2476; e-mail: james.eberhardt@ee.doe.gov 

Field Technical Manager: Philip S. Sklad 

(865) 574-5069; fax: (865) 574-6098; e-mail:skladps@ornl.gov 

Contractor: Oak Ridge National Laboratory 

Contract No.: DE-AC05-00OR22725 

Objectives 

• Develop wrought magnesium (Mg) alloys with better formability than current Mg alloys. 

• Investigate new processing techniques for cost reduction and formability improvement. 

• Contribute to basic understanding of deformation, processing, and alloy behavior for this lightweight metal. 

Approach 

• Explore new processing schemes for Mg alloy sheets, evaluate the practical formability of new alloys and alloys subjected to new processes, determine deformation mechanisms using experimentation and simulation. 

Accomplishments 

• Presented and published roll bonding paper. 

• Published paper on deep drawing of sheet Mg. 

• Presented and published paper on alloying for texture modification; it demonstrates new path to enhanced 

formability. 

• Initiated rolling studies of new set of alloys cast by MEL. 

Future Direction 

• Interact with an industrial partner to advance commercialization of infrared processing for cost reduction. 

• Define further correlations of microstructure, mechanisms, and anisotropy. In particular, define the kinetics and 

crystallography of recrystallization in Mg alloys. 

• Further explore superplastic-type behavior observed at warm forming temperatures. 

39
FY 2006 Progress Report 

Introduction 

Magnesium, with a density that is 66% of that of 

aluminum, has obvious advantages for use as a 

structural material in the transportation industry. 

Magnesium castings are steadily growing in use in 

automobiles and trucks. However, wrought applications are held back by the high cost of sheet due to 

the large number of processing steps. This project is 

currently focusing on processing techniques and issues to reduce the cost of sheet Mg. In addition, we 

continue to investigate means to improve the formability of wrought Mg alloys. 

During this period, we have finished and submitted several publications based on work largely performed last year with details presented earlier. 

Rolling Studies of New Alloy Compositions Cast at MEL for Cost Reduction 

Wrought alloy costs have several components, 

including starting alloy composition and original 

ingot casting procedures. A collaborative effort with 

MEL is aimed at addressing this issue in conjunction 

with tailoring an alloy for a continuous cast and rolling line. Ninety small slabs were cast at MEL for 

rolling and characterization studies at Oak Ridge 

National Laboratory (ORNL) (see Figure 1). Both 

normal rolling schedules and schedules amenable to 

an infrared reheat continuous process are being investigated. Previous work at ORNL has shown that 

infrared reheat yields similar results to current commercial practice suggesting that, in conjunction with 

twin roll casting, it presents a possibility for a continuous production line. 

Figure 1. An initial rolling attempt at a new composition 

alloy. 

Deformation Behavior of a Twin-roll-cast 

Alloy 

Mechanical properties of a twin-roll-cast 

Mg6Zn0.5Mn0.3Cu0.02Zr with a grain size of ~15 μm 

were measured at temperatures from 423 to 523 K. 

Constant strain rate, strain rate change, and temperature change tests (strain rates from 5×10−7

 to 

High Strength Weight Reduction Materials 

3.3×10−3/ s−1) were employed to characterize the 

plastic flow behavior of the alloy (Figure 2). It was 

observed that, at low strain rates and high temperatures, the alloy followed a power law behavior; but 

the power law breaks down at high strain rates and 

low temperatures. Transient behavior was also characterized to identify the dominant deformation 

mechanism in the alloy. 

Figure 2. Strain rate vs steady state flow stress for a twinroll-cast Mg alloy. 

Roll Bonding 

The process of accumulative roll bonding 

(ARB) has been applied to wrought Mg alloy AZ31. 

Accumulative roll bonding results in severe plastic 

deformation of sheet material. In AZ31, this process 

leads to the refinement of the microstructure by dynamic recrystallization. The refinement of the microstructure results in an improvement in 

mechanical properties with the potential for producing materials for superplastic forming. In this work, 

AZ31 was bonded with up to 64 layers and the potential of roll cladding with aluminum was investigated. The development of microstructure and the 

mechanical properties are reported in this document. 

The conclusions of the paper: 

• Successful ARB of AZ31-H24 sheets was 

achieved by rolling at 400oC. 

• It was possible to roll Mg clad with aluminum 

alloy 6061. 

40
High Strength Weight Reduction Materials 

• ARB results in significant grain refinement by 

dynamic recrystallization. This refinement of the 

microstructure leads to significant improvements 

in the high-temperature ductility (Table 1) and 

indicates the possibility for this processing route 

to be used to produce sheet amenable to superplastic forming. 

FY 2006 Progress Report 

and Prangnell.

1

 Samples were deformed in plane 

strain compression to prescribed strain levels at temperatures high enough to prevent cracking and then 

were recrystallized. The deformation texture was 

very similar to that observed after plane strain compression (e.g., rolling) of other Mg alloys, such as 

common AZ31 (Figure 5). However, the texture can 

radically evolve during recrystallization. Annealing 

Table 1. Mechanical properties of roll-bonded AZ31 

Grain RMT 400EC 

Condition 

size 

(μ m) Elongation 

(%) 

UTS 

(MPa) 

Elongation 

(%) 

UTS 

(MPa) 

AZ31 H24 ~10 12.2 294.5 37.9 20.5 

ARB1 64 layers Min. 1 9.0 234.0 137.9 22.6 

ARB2 A1’-Mg’-‘A1 Mg-5 15.2 199.5 64.8 29.5 

ARB3 AZ31-70% 5 21.8 288.0 126.9 21.5 

Unlike aluminum alloys, the texture of ARB Mg 

alloy AZ31 is very similar to the texture of conventional hot-rolled AZ31. 

Deep Drawing 

The sheet formability of current Mg alloys at 

ambient temperatures is poor; however, the formability at moderately elevated temperatures can be 

excellent. Cylindrical cup drawing tests are used to 

compare the warm forming characteristics of conventional alloy AZ31B with those of alloys containing lithium or yttrium solid solutions. While both 

types of experimental alloy may have better room 

temperature ductility (e.g., εf ~ 25–30%) than 

AZ31B, only the lithium alloy has comparable or 

better deep drawing capacity. The results are discussed in terms of the sheet anisotropy, and attention 

is drawn to the fact that Mg alloys exhibit an unusual bending response because of mechanical twinning. The analysis has further shown that the 

strength of the Mg sheet is sufficient to pull the 

sheet into the die; however, difficulties reside in the 

bend/unbend and ironing operations that occur at the 

radius of the top edge of the die. (See Figures 3 

and 4.) 

Alloy Development 

A Mg alloy modeled after the commercial alloy 

WE54 was thermomechanically processed and characterized to elucidate the mechanisms by which its 

texture can be randomized, as reported by Ball 

above the precipitate solvus temperature results in a 

nearly random recrystallization texture where the 

initially strong basal texture is completely eliminated. Annealing below the solvus temperature produces a recrystallized texture similar to that of 

conventional Mg alloys, i.e., a weakened basal texture. If this same material is then annealed above the 

solvus to promote grain growth, a randomized texture evolves, similar to that of samples simply recrystallized above the solvus. It is hypothesized that 

a sort of preferred growth mechanism, whereby 

grains with random orientation grow faster than 

those with the original basal texture, contributes significantly to the texture randomization observed in 

wrought Mg alloys containing rare earth elements. 

Mechanical Modeling of Forming 

O-temper AZ31B Mg sheet was deformed under 

large-strain tension/compression. The effects of 

twinning and untwinning, which are responsible for 

asymmetric yield and hardening evolution, were 

revealed by metallography, acoustic emission (AE), 

and texture measurements (Figure 6). Predominant 

deformation mechanisms were revealed: basal slip 

for initial tension, twinning for initial compression, 

and untwinning for tension following compression. 

A working model consistent with microscopic and 

macroscopic behavior was constructed. 

41
FY 2006 Progress Report High Strength Weight Reduction Materials 

Figure 3. Effect of temperature on drawing operations. 

Figure 4. Rate effects on drawing on Mg alloys. 

500 

450 

(Times-Random) x100 400 

350 

300 

250 

200 

150 

100 

50 

0 

0 10 20 30 40 50 60 70 80 90 

Tilt from ND (o) 

a 

c 

b RXXed below solvus sample 2 (ε = 0.64, v = 5 mm/s) 

re-annealed above solvus sample 2 

(ε = 0.64, v = 5 mm/s) 

RXXed above solvus sample 6

 (ε = 1.15, v = 20 mm/s) 

Figure 5. Plot of basal intensity vs tilt showing how texture was eliminated during 

recrystallization of the Mg5Y3Nd1Zr alloy. 

Reference Mg-6Zn-0.5Mn-0.3Cu-0.02Zr Alloy at Elevated 

Temperatures” submitted to Script Met. 1. E. A. Ball and P. B. Prangnell, Scripta Metall. A. L. Bowles and J. A. Horton, “Accumulative Mater. 31, 111–116 (1994) . Roll Bonding of Magnesium Alloy AZ31,” CIMTMS Symposium on Magnesium Technology in the New Publications and Presentations Global Age, Montreal, Canada, October 1–4, 2006. 

S. X. Song, J. A. Horton, N. J. Kim, and T. G. S. R. Agnew, J. W. Senn, and J. A. Horton, “Mg 

Nieh, “Deformation Behavior of a Twin-roll-cast Sheet Metal Forming: Lessons Learned from Deep 

42
High Strength Weight Reduction Materials FY 2006 Progress Report 

Figure 6. Large strain tension/compression curves of 

AZ31 used for forming/springback models. 

Drawing Li and Y Solid-Solution Alloys,” JOM 

58(5), 62–69 (2006). 

S. R. Agnew, D. W. Brown, and C. N. Tomé, 

“Validating a Polycrystal Model for the Elastoplastic Response of Magnesium Alloy AZ31 Using 

In-situ Neutron Diffraction,” Acta Materialia 54, 

4841–4852 (2006). 

J. W. Senn and S. R. Agnew, “Texture Randomization during Thermomechanical Processing of 

a Mg-Y-Nd Alloy,” Proceedings of the Magnesium 

Technology in the Global Age Symposium, Montreal, 

Canada, October 2006. 

J. Bohlen, M. R. Nürnberg, J. W. Senn, D. 

Letzig, and S. R. Agnew, “The Texture and Anisotropy of Magnesium-Zinc-Rare Earth Alloy Sheets,” 

Acta Mater., in press. 

X. Y. Lou, M. Li, R. K. Boger, S. R. Agnew, 

and R. H. Wagoner, “Hardening Evolution of 

AZ31B Mg Sheet,” Inter. J. Plasticity 23, 44–86.6 

(2007). 

Previously Reported Publications and 

Presentations 

A. L. Bowles, Q. Han, J. A. Horton, “Castability 

of Magnesium Alloys” pp. 99–104 in Magnesium 

Technology 2005, eds. N. R. Nelameggham, H. I. 

Kaplan, and B. R. Powell, TMS, 2005. 

J. A. Horton, C. A. Blue, T. Muth, A. L. 

Bowles, S. R. Agnew, “Infrared Processing of Magnesium Wrought Alloys” 155–158 in Magnesium 

Technology 2005, eds. N. R. Neelameggham, H. I. 

Kaplan, and B. R. Powell, TMS, 2005. 

S. R. Agnew and O. Duygulu, “Plastic Anisotropy and the Role of Non-basal Slip in Magnesium 

Alloy AZ31,” Inter. J. Plasticity 21, 1161–1193 

(2005). 

S. R. Agnew, D. W. Brown, and C. N. Tomé, 

“Validating a Polycrystal Model for the ElastoPlastic Response of Magnesium Alloy AZ31 Using 

In-situ Neutron Diffraction,” submitted to Acta Mater. 

S. R. Agnew, P. Mehrotra, and T. M. Lillo, 

“Equal Channel Angular Extrusion of a Heat Treatable Magnesium Alloy,” Scripta Mater., in revision. 

S. R. Agnew, P. Mehrotra, T. M. Lillo, G. M. 

Stoica, and P. K. Liaw, “Crystallographic Texture 

Evolution of Three Wrought Magnesium Alloys 

during Equal Channel Angular Extrusion,” Mat. Sci. 

Eng. A 408, 72–78 (2005). 

S. R. Agnew, P. Mehrotra, T. M. Lillo, G. M. 

Stoica, and P. K. Liaw, “Texture Evolution of Five 

Wrought Magnesium Alloys during Route A Equal 

Channel Angular Extrusion: Experiments and Simulations” Acta Mater. 53, 3135–3146 (2005). 

D. W. Brown, S. R. Agnew, M. A. M. Bourke, 

T. M. Holden, S. C. Vogel and C. N. Tomé, “Internal Strain and Texture Evolution during Deformation Twinning in Magnesium,” Mat. Sci. Eng. A. 

399, 1–12 (2005). 

D. W. Brown, S. R. Agnew, S. Abeln, W. Blumenthal, M. A. M. Burke, M. Mataya, C. N. Tome, 

and S. C. Vogel, “The Role of Texture, Temperature 

43
FY 2006 Progress Report 

and Strain-rate in the Activity of Deformation Twinning,” Mat. Sci. Eng. A 495–497, 1037–1042 (2005). 

A. Jain and S. R. Agnew, “Modeling the Temperature Dependent Effect of Twinning on Mechanical Behavior of Magnesium,” in Tenth Inter. Symp. 

on the Physics of Materials, Prague, Czech Republic, August 2005 (in press, Mat. Sci. Eng. A) 

S. R. Agnew, O. Duygulu, and A. Jain, “Using 

Polycrystal Plasticity Simulation to Explore the Behavior of Wrought Magnesium Alloys,” presented at 

Plasticity 2005, Kauai, January 2005. 

S. R. Agnew, O. Duygulu, and A. Jain, “Developing Mechanism-Based Models of the Mechanical 

Behavior of Wrought Magnesium Alloys,” presented 

at the University of Michigan Department of MSE 

Seminar Series, Ann Arbor, Michigan, February 

2005. 

High Strength Weight Reduction Materials 

O. Duygulu and S.R. Agnew, “Grain Size Effects on Deformation Mechanism Activity in Magnesium Alloy AZ31B,” presented at the Magnesium 

Technology 2005 Symposium, TMS Annual Meeting, San Francisco, February 2005. 

S. R. Agnew, D. W. Brown, and C. N. Tomé, 

“Neutron Diffraction Study of Deformation Mechanisms in Magnesium Alloy AZ31B,” presented at 

the Symposium on Neutron Diffraction Characterization of Mechanical Behavior, TMS Annual Meeting, San Francisco, February 2005. 

44
High Strength Weight Reduction Materials FY 2006 Progress Report 

D. Alloy Design and Thermomechanical Processing of a Beta Titanium Alloy for 

a Heavy Vehicle Application 

Principal Investigator: Craig A. Blue 

Oak Ridge National Laboratory 

Oak Ridge, TN 37831-8063 

(865) 574-4351; fax: (865) 574-4357; e-mail: blueca@ornl.gov 

Field Project Manager: William H. Peter 

Oak Ridge National Laboratory 

Oak Ridge, TN 37831-8063 

(865) 241-8113; fax: (865) 574-4357; e-mail: peterwh@ornl.gov 

Chief Scientist: James J. Eberhardt 

(202) 586-9837; fax: (202) 587-2476; e-mail: james.eberhardt@ee.doe.gov 

Field Technical Manager: Philip S. Sklad 

(865) 574-5069; fax:(865) 574-6098; e-mail: skladps@ornl.gov 

Participants: 

Edwin H. Kraft, EHK Technologies 

Evan K. Ohriner, Oak Ridge National Laboratory 

David C. Harper, Oak Ridge National Laboratory 

Jim O. Kiggins, Oak Ridge National Laboratory 

Donny McInturff, Oak Ridge National Laboratory 

Lance Jacobsen, International Titanium Powder, L.L.C. 

Bill Ernst, International Titanium Powder, L.L.C. 

Dariusz Kogut, International Titanium Powder, L.L.C. 

DOE CRADA Partner−Producer of Heavy Vehicle Components 

Contractor: Oak Ridge National Laboratory 

Contract No.: DE-AC05-00OR22725 

Objectives 

• Develop a “low cost” beta titanium (Ti) alloy to reduce the weight of a heavy vehicle component, using a lowcost powder developed by the Defense Advanced Research Projects Agency (DARPA) Titanium Initiative or 

another low-cost Ti powder. 

• Investigate blending procedures for elemental or master alloy additions and homogenization (specifically for 

solid state processing). 

• Develop processing parameters for solution treating, rolling, and aging of the new Ti alloy. 

• Evaluate microstructure, mechanical behavior, and chemistry of arc-melt and drop castings of new alloy to 

compare with existing, expensive beta Ti compositions. 

• Demonstrate the ability to extrude rod using powder Ti precursor. 

• Demonstrate “low-cost” production of beta Ti alloy component. 

45
FY 2006 Progress Report High Strength Weight Reduction Materials 

Approach 

• Collaborate with the rest of the DOE team to develop new alloy compositions specifically geared toward DOE 

heavy vehicle/truck components. Mechanical properties, similar microstructures to existing beta alloys, and 

cost are the main parameters that will determine the success of the new alloy. 

• Cast the new compositions along with conventional compositions compare cast, solution-treated, cold-worked, 

and aged microstructures. Tensile properties, microstructures, and cost will determine which composition will 

ultimately be extruded. 

• After down-selecting to one composition, extrude blended powders to full consolidation and product form. 

• Develop the necessary thermomechanical processing techniques for the desired properties required for end use. 

• Ultimately, produce component from new alloy. 

Accomplishments 

• Extruded CP Ti, Ti-6Al-4V, and other Ti alloys from low-cost ITP powder with comparable mechanical 

properties, microstructures, and chemistries to ASTM standards. 

• Selected two beta Ti alloy compositions for casting. 

• Made several arc-melt castings of a new alloy composition and a conventional beta composition for 

comparison. Performed differential thermal analysis and chemical analysis on the castings. 

• Initiated study of preliminary elemental blending of powders and cold consolidation of powder to increase 

apparent density. 

• Redesigned component to take into consideration Ti beta alloy mechanical properties. CRADA partner initiated 

studying necessary cold work processing steps for optimum mechanical properties using conventional beta 

alloy. 

Future Direction 

• Perform monotonic tensile testing and chemistry analysis on additional castings of new alloy compositions. 

• Continue to improve blending techniques while limiting oxygen pickup. 

• Extrude new Ti beta alloy powder compositions. 

• Optimize thermomechanical processing technique required for component. 

• Produce beta Ti alloy component. 

• Conduct laboratory and vehicle testing of Ti component. 

Introduction 

The aircraft industry is currently the single largest market for Ti and Ti alloy products primarily 

because of the exceptional strength-to-weight ratio, 

elevated temperature performance, and corrosion 

resistance. The strong dependence of Ti on the aerospace industry has caused Ti production to be very 

cyclic. Currently, Ti alloy sheet costs are around 

$40–50/lb, and lead times for product delivery from 

time of order can be as long as 18 months. Titanium 

usage outside the aerospace industry has been limited by its higher cost relative to competing materials, primarily aluminum alloys and steels. Although 

Ti possesses an attractive set of properties— 

including high specific strength, corrosion resistance, damage tolerance, low Young’s modulus, 

high strength-to-weight ratio, and biocompatibility—its cost limits its applications to selected markets. Many automotive systems would benefit from 

the use of Ti products. Automotive exhaust systems 

could save as much as 50% of their current weight 

by integrating Ti parts.1,2,3

 Titanium valves and 

valve springs, connecting rods, suspension springs, 

wheels, drive shafts, underbody panels, side impact 

46
High Strength Weight Reduction Materials 

bars, and half shafts are just some of the automotive 

applications that could benefit from the use of Ti. 

A recent market comparison for the use of engineering materials shows that steel is the most widely 

used material, with 725 million tons used each year.4 

Aluminum, stainless steel, and copper are used at a 

rate of 20 million tons, 14.5 million tons, and 11 

million tons, respectively.

4

 Titanium is much less 

widely used at 0.04 million tons per year.4

 Government studies have shown that Ti enjoys only 1/20th 

of its potential usage.5 

Technical problems preventing the integration of 

Ti, apart from the cost of manufacturing, include 

wear resistance, a lower modulus than steel, and machining difficulties. Wear resistance can be addressed by coatings, reinforcements, and 

compositing, as has been shown in preliminary work 

performed by Blau et al.6

 Low modulus is a benefit 

in some applications; when required, modulus can 

be increased by reinforcing the part with a second 

material. The machining difficulties can be reduced 

by the production of near-net-shape parts. The major 

problem is that Ti costs substantially more than 

competing materials. 

Although the cost of Ti ore is significantly more 

than that of other competing materials, the difference in the cost of sheet, plate, and most other Ti 

and Ti alloy product forms is a result of secondary 

processing or milling operations (Figure 1). For example, processing of Ti from ingot to 1-in.plate accounts for 47% of the total cost of the material.7,8 

Thinner plate and sheet gauge sections increase the 

production costs, the amount of production waste, 

and the percentage of total production costs. 

Current melting techniques include vacuum arc 

remelting (VAR), electron beam melting, and 

plasma arc single-melt processes. VAR is used to 

consolidate feedstocks, improve homogeneity, and 

refine grain size. To VAR-cast an ingot, a dc arc is 

struck between a Ti electrode and a base plate. The 

intense heat melts the electrode, and the Ti is cast 

into an ingot. VAR requires 20 hours and large-scale 

equipment. 

Electron beam and plasma arc single-melt processes have the advantage over VAR in increased use 

of low-cost scrap and improved removal of highdensity and low-density inclusions. In these two 

FY 2006 Progress Report 

Figure 1. Breakdown of cost in the production of 1­

in. CP Ti plate. 

processes, ingots are cast directly from melted feed 

sponge, alloy additions and scrap. But the cast ingot 

still requires post-processing, which accounts for 

47% of the total cost.7,8 

Although much time and effort has been invested in the development of these processing techniques, the cost of Ti still remains high because of 

yield losses and the cost of raw materials and postprocessing. In an ongoing effort to reduce the cost of 

Ti and Ti alloy production, Oak Ridge National 

Laboratory (ORNL) has investigated new solid state 

and melt technologies in producing plate, sheet, and 

rod. Vacuum hot pressing (VHP), high-density infrared (HDI) processing, roll compaction, and extrusions have been explored at ORNL (Figure 2).3,9 

These methods used DARPA Titanium Initiative 

“low-cost” powders produced by the Armstrong 

method and provided by International Titanium 

Powder (ITP). ITP has produced CP Ti and Ti alloys 

in its pilot processing facility and has initiated the 

development of a full-scale production plant to produce CP Ti and Ti alloy powders. Consolidation 

studies are required to produce consolidated components, parts, and material. 

This annual progress report discusses recent 

studies to develop a low-cost beta Ti alloy for a 

heavy vehicle component. This work is performed 

with a vehicle component manufacturer that currently produces the targeted component from steel. 

The producer has seen the value added by potentially replacing the component with a Ti alloy if 

47
FY 2006 Progress Report 

Figure 2. Various product forms of consolidated ITP 

CP Ti and Ti alloy powder resulting from manufacturing process development at ORNL. 

production costs can be minimized. ORNL is collaborating with this company to determine a new 

alloy composition, lower in price than current aerospace Ti alloy compositions, that is geared towardthe mechanical properties of the component. Arcmelted drop castings are produced of the new compositions to evaluate the mechanical characteristics 

compared with existing expensive compositions. 

ORNL will then develop a material processing technology that successfully consolidates blended powder, produces the necessary characteristics, and 

reduces the cost of current Ti production mill processing. 

Experimental Procedure 

The present scope of work was performed 

within four tasks: 1, component design; 2, economical modeling; 3, alloy design and verification; and 4, 

process development. Each task was designed with 

the ultimate goal of producing a beta Ti alloy component that reduced the weight of a heavy truck 

component. For ease of discussion, the experimental 

procedure and the results will be subdivided into the 

four tasks. 

Task 1: Component Design 

The identity of the component selected for this 

CRADA is proprietary. However, certain mechanical/structural considerations had to be taken into 

account to redesign the component so that a Ti alloy 

High Strength Weight Reduction Materials 

could be successfully implemented. Component design was performed by the CRADA partner, with 

ORNL providing information on the potential processing routes and thermomechanical history that 

could influence the resulting product form and component design. 

Task 2: Economic Model 

A preliminary economic model was constructed 

based on a “greenfield” production approach. Specific economics were taken into account include labor, price of the Ti powder, price of elemental 

alloying powders, thermomechanical processing 

costs, machining, capital costs, and post-aging processing. The model has been modified as the processing development matures. The resulting economic 

model will be comprehensive. 

Task 3: Alloy Design and Verification 

The selected component would best use an alloy 

with the beta (body centered cubic) crystal structure. 

Beta alloys are normally composed of beta grains 

with precipitates of the alpha (hexagonal closepacked) phase within the beta grains and at grain 

boundaries. The application of the selected component will operate at ambient temperatures, so there is 

no need to provide high-temperature strength, creep 

resistance, or enhanced oxidation resistance; but Ti 

corrosion resistance is a benefit. One priority objective of the alloy design is the achievement of a comparable low-cost alloy. Many of the conventional 

commercial alloys contain significant levels of expensive alloy additions. Since the CRADA project 

uses powder processing rather than melt metallurgy, 

issues of melt segregation of elements, such as iron, 

are not of concern. Therefore, lower-cost beta stabilizing elements can be used in the design of the 

composition for the heavy vehicle component. 

A method of alloy design was used that takes 

advantage of established tendencies of various elements to stabilize the alpha or beta phases in comparison with the benchmark elements molybdenum 

(Mo) (beta stabilizer) and aluminum (Al) (alpha stabilizer); these tendencies are known as the Mo 

equivalence or Al equivalence of each element. A 

beta stability index for an alloy is then calculated as 

the sum of the Mo equivalences of all its beta stabilizing constituents, minus the sum of the Al equiva48
High Strength Weight Reduction Materials 

lences of all its alpha stabilizing constituents; this is 

the Mo–Al equivalence. This index was calculated 

for many commercial beta alloys and for numerous 

candidate compositions that would contain low levels of expensive elements and could use inexpensive 

master alloy powders. 

One primary and two secondary compositions 

were selected for further investigation. Transformation diagrams were determined by computer simulation for these three compositions. The primary 

composition and a conventional beta Ti composition 

were then arc-melted and drop cast to provide material for preliminary processing. The intense heat of 

the arc was used to melt the Ti and other constituent 

elements of the alloy. Elements were thoroughly 

mixed in a copper crucible and turned over four to 

five times between arc-melting operations to improve homogeneity. Once mixed, the alloy button 

was cast in a water-cooled 1⁄2×1⁄2×3 in. copper mold. 

Chemical analysis was performed on the Ti alloy 

ingots, inert gas fusion was used to determine oxygen and nitrogen weight percentages (ASTM E 

1409-05), and metal elements were analyzed using 

direct current plasma emission spectroscopy (ASTM 

E 1097-03). Micrographs were made of the as-cast 

microstructure. Small samples (20 to 50 mg) were 

cut from the castings to perform differential thermal 

analysis (DTA) measurements using a SDT Q600 

from TA Instruments. A heating rate of 20°C/min 

was used, and the sample was allowed to go a full 

cycle. A CP Ti sample cast from low-cost ITP Ti 

was analyzed for comparison. DTA analysis was 

performed to identify the beta transus. 

Task 4: Process Development 

A solid state or powder metallurgy processing 

technique that takes advantage of the new low-cost 

Ti powders is the objective. The blending of elemental or master alloy powders to form the desired alloy, 

extrusion parameters and methodology, solution 

treating temperatures and time intervals, rolling conditions, aging temperature and time intervals, and 

post-aging conditioning are all thermomechanical 

variables that must be developed and identified for 

the realization of the heavy vehicle Ti alloy component. However, identifying these parameters on extruded samples of blended powders initially would 

be costly and time-prohibitive. Therefore, prelimiFY 2006 Progress Report 

nary solution treating (ST), cold rolling (CR), and 

aging (A) processes are being performed on drop 

castings to evaluate the candidate alloys in the required processed state, since cast microstructures 

will not provide the necessary mechanical properties 

for the application. Once an alloy composition exhibits the required microstructures, economics, and 

mechanical properties, extrusions of the alloy will be 

initiated and thermomechanical processing parameters will be refined for the extruded condition. Process development is being performed in four 

individual efforts to rapidly develop the processing 

technology. The four stages include blending of 

elemental powders (performed at ORNL); extrusion 

of low-cost Ti powders (performed at ORNL); development of ST, CR, and A (performed at ORNL); 

and post-aging conditioning of the component (performed at the CRADA partner’s facilities). Procedures are identified for blending of elemental 

powders and development of ST, CR, and A procedures in the following paragraphs. The extrusion 

procedures for low-cost Ti powder were discussed in 

the FY 2005 Annual Progress Report.9

 Post-aging 

conditioning is proprietary to the CRADA partner 

and will not be discussed. 

Blending of elemental powders 

The blending of powders to establish the alloy 

was performed using ITP Ti powder synthesized by 

the Armstrong process and elemental alloying powders. The alloying powders were not of high purity 

for the first trials performed in this fiscal year. The 

initial objective was to blend the powders using attrition methods and observe the degree of mixing or 

homogenization that occurs. The secondary objective will be to minimize the level of oxygen pickup 

during blending operations. Once homogenization 

has been verified, high-purity powders will be used. 

Two batches of Ti powder with elemental alloying 

powders were milled, around 128 g for each batch. 

One batch was dry milled, and the other was wet 

milled in ethanol to see if one method enabled 

greater proficiency than the other. A zirconia media 

at a ball-to-powder weight ratio of 14 to 1 was used 

to blend the alloy and attrite the ITP Ti powder particles. Milling was performed for 4 hours. The drymilled bottle was pumped out and backfilled two 

times with argon. The milling rate was around 

49
FY 2006 Progress Report High Strength Weight Reduction Materials 

100 rps in a 1-L high-density polyethylene container. The material milled in ethanol was dried at 

60°C overnight and then oblique blended for 1 h to 

mix up the powder after settling during the drying 

process. The blended powders were then compressed 

into 1.7-cm-diam disks, 3–6 mm thick, at a pressure 

of 20,000 psi in air. In a similar cold compression 

process, ITP CP Ti powder, without alloying elements and without milling operations, was also 

pressed at various pressures to assess compaction 

versus pressure trends for the ITP powder morphology. The results will later help to develop the pressures required to consolidate powder into cans for 

extrusion. (The CP Ti powder will have lower densities per a given pressure than a milled and blended 

powder, but it will provide minimal pressure requirements to obtain desired cross sections in the 

extruded condition). 

Scanning electron microscopy (SEM) and energy-dispersive X-ray (EDX) mapping were performed on the compressed alloy samples to observe 

homogeneity. Chemical analysis was also performed 

to assess the accuracy of the chemical composition; 

inert gas fusion was used to determine oxygen and 

nitrogen weight percentages (ASTM E 1409-05); 

and metal elements were analyzed using direct current plasma emission spectroscopy (ASTM E 1097­

03). 

Development of ST, CR, and A procedures 

Development of ST, CR, and A procedures is 

critical for the resulting mechanical properties of a 

Ti alloy component. The primary developed alloy 

and the conventional beta Ti alloy were both studied. ST of the as-drop-cast ingots was performed in a 

small circular halogen/T3 infrared furnace for 0.5 to 

1 h at temperatures above the beta transus in ultrahigh-purity argon atmosphere. CR was performed on 

a 4-high WC roll mill. Multiple passes were performed on the ST samples; 3–5% reductions were 

taken on every pass. Total CR reductions were up to 

50%. Aging was performed in a tungsten-element, 

vacuum furnace at moderate temperatures for 4 to 30 

h. Macro and micro hardness values were measured 

throughout the processing of both compositions. 

Hardness values and microstructures were compared. 

Results and Discussion 

Task 1: Component Design 

Initial design of this component in Ti was completed. It was selected to replace a steel vehicle 

component for a weight savings of 40%, including 

allowance for the geometric changes required to 

provide equal mechanical performance. The component is of a bar shape amenable to extrusion, which 

was selected as the primary process, as well as alternative process methods. 

Task 2: Economic Model 

A preliminary greenfield model was produced 

for production of the Ti alloy component. However, 

after thorough discussion among all parties at the 

annual review meeting, the approach was modified 

to focus on the use of existing commercial operations wherever possible. Restructuring of the model 

to accommodate this change began, and an effort 

also began to identify potential facilities to use in 

volume production. A synergistic approach is required for the realization of the Ti component and 

the economic model, since economics are of primary 

importance, new production capabilities are continuing to evolve, and Ti solid state processing is a new 

and rapidly expanding field. The economic model 

will continue to develop as process development 

continues. 

Task 3: Alloy Design and Verification 

As previously indicated, a primary low-cost 

composition was selected for initial development. 

The low-cost composition and the conventional beta 

Ti alloy were first cast into buttons and then successfully drop cast into square cross-sectioned ingots. At least three ingots were produced for each 

composition. Figure 3 shows a button and drop casting of the new composition, a micrograph of the new 

composition, and a micrograph of the conventional 

composition. 

50
High Strength Weight Reduction Materials 

Figure 3. Composite image of (a) button and drop 

cast ingot, (b) optical micrograph of the new composition, and (c) optical micrograph of conventional beta 

Ti alloy processed in a similar manner. 

The drop cast ingot and button of the conventional beta alloy and the new alloy were similar in 

color and appearance. The micrographs exhibited 

equiaxed grains, with the new composition having a 

similar appearance to the conventional composition. 

The chemical analysis results were very good for 

preliminary trials. The conventional alloy contained 

one alloying element that was off by 1 wt %, and the 

Ti composition was slightly high. The developed 

composition was closer to target with alloying elements within 0.3 wt =% of target values. Interstitial 

elements (e.g., oxygen, nitrogen) were very low for 

both compositions. The hardness of the as-cast ingots was the only significant difference observed 

between the two alloys; the hardness value of the 

new alloy was approximately 40% greater than that 

of the conventional alloy. The increase in hardness 

seemed to be based solely on the metal alloying additions, as the oxygen and nitrogen content of both 

sets of ingots was extremely low. Future castings 

will use cold-pressed ITP CP Ti powder pellets, as 

initial castings with pure Ti from an outside source 

had oxygen values much lower than required. Additionally, a third composition that may eliminate requirements for a powder blending procedure will be 

investigated. 

Task 4: Process Development 

Blending of elemental powders 

Initial blending efforts focused on the required 

procedures to homogenize the powders. Preliminary 

FY 2006 Progress Report 

chemical homogenization results were promising; 

SEM and EDX results exhibited relatively good uniformity to around 50 to 100 μm. Several of the initial powder particles were larger than optimal; this is 

the main reason chemical uniformity was not observed below 50 μm. However, fine alloying element powder particles exhibited good blending at 

even very low scales, as can be seen in Figure 4. The 

chemical analysis results were within 0.6 wt % for 

each alloying element. Oxygen and nitrogen values 

were high in the partially consolidated blended 

powders; this may be a result of initially high oxygen levels in the alloying powders used and oxidation during the ball milling process. Values of 

zirconium, from the zirconia ball media, were below 

traceable values in the chemical analysis results. 

Figure 4. SEM of partially consolidated blended alloy. Fine alloy powder particles were well distributed 

through ITP Ti particles. 

Extrusion procedures 

Extrusions for various sizes, shapes, pressures, 

and compositions have been previously reported. 

Fully dense, as-extruded samples have exhibited 

mechanical, chemical, and metallurgical properties 

similar to those of wrought Ti products. Photos, micrographs, and the resulting mechanical properties 

are provided in Figure 5. 

51
FY 2006 Progress Report High Strength Weight Reduction Materials 

Figure 5. Photos of CP Ti after removing can and 

machined into round tensile bars, with mechanical 

results and micrographs of extruded bars and flat 

stock. 

The next major step in the extrusion process development is the extrusion of the blended alloy 

powder; this cannot be performed until the blending 

technique is optimized. However, developing methodologies to increase the tap density of the low-cost 

Ti powder is an issue that directly influences the 

possible size of resulting extrusions and is presently 

under investigation. In FY 2006, cold press consolidation pressures were compared with the resulting 

apparent densities. Previously, ITP powder was cold 

pressed in the extrusion can to ~20% of theoretical 

density. Higher pressures are required to extrude 

larger cross sections. Figure 6 is a graph illustrating 

low pressures and the resulting densities; this information will be used in the extrusion of the beta Ti 

alloys for component fabrication. 

Figure 6. Graph illustrating the resulting density 

when various compaction pressures are applied to ITP 

Ti powder in a steel die with lubricated walls. 

Development of ST, CR, and A procedures 

Castings of the new alloys have made it possible 

to begin the thermomechanical process development 

required for the desired mechanical properties prior 

to extrusion of the blended powders. Therefore, ST, 

CR, and A procedures were initiated. However, the 

main intent of the preliminary thermomechanical 

processing at this stage is to allow for proper alloy 

selection. Preliminary work was performed mainly 

on the arc melt castings of the conventional alloy 

and compared with VAR-melted samples prepared 

from an outside commercial source using conventional techniques. The new alloy will be processed 

similarly to the conventional material using the beta 

transus, acquired from DTA results, as a reference 

temperature. CR of the cast, conventional alloy was 

performed; a total reduction of 50% after multiple 

passes was possible. The cast and ST new alloy will 

be compared with this amount of cold work. However, the new composition is harder than the conventional alloy, and the total CR reduction may not need 

to be at 50%. Figure 7 shows the conventional alloy 

during and after processing with microstructures of 

the cast material compared with conventional VAR 

material. 

Figure 7. Image of the ORNL conventional alloy ascast, ST, and CR/A (left) with microstructures of the 

ST and ST+CR/A conditions (center). For comparison, micrographs of a commercially produced VAR 

sample are included (right). 

Post-aging procedures and development 

of the component 

The CRADA partner performed preliminary final processing on commercial beta C plate material 

52
High Strength Weight Reduction Materials 

to begin understanding the effect of this processing 

on properties. The details of this processing are proprietary. However, the effort resulted in a basic understanding of the relative effect of several process 

parameters on some critical material properties. 

These data will be useful in specifying these process 

parameters for powder-processed material of the 

finally selected composition. 

The CRADA partner also performed aging studies to begin an understanding of the effect of aging 

time and temperature on properties. Rockwell C 

hardness was measured after various aging times at 

four aging temperatures. Yield strength and hardness 

were then measured for two of the most promising 

aging conditions. 

All partners in this project have also participated 

in gathering data and literature on the processing, 

heat treatment, and properties of various beta alloys 

for use in designing thermomechanical processing 

and heat treatment of the powder processed alloys. 

Summary 

The following are highlights of progress in 

FY 2006. 

1. Extrusion of CP Ti, Ti-6Al-4V, and other Ti 

alloy was performed in the production of rods, 

bars, and flat stock. 

2. A primary alloy composition was developed. 

3. An economic model was initiated. 

4. Castings of a primary low-cost composition and 

a conventional alloy for comparison were fabricated. Subsequent ST, CR, and A process development was initiated with promising preliminary 

results. 

5. Blending technique development was initiated. 

6. Post-aging procedures are being developed at 

the CRADA partner’s organization. 

Conclusions 

Although Ti and Ti-6Al-4V have attractive 

properties that would benefit a great number of engineering applications, their use is limited by the 

cost of the raw material and the processing necessary to make a usable product. But with recent advancements in low-cost powder production at ITP, 

FY 2006 Progress Report 

in conjunction with processing techniques developed 

at ORNL in collaboration with industry, the cost of 

Ti and Ti alloy products could be drastically reduced. Previous extrusions have proved the feasibility of using low-cost powder in new, innovative 

markets. The preliminary study of CP-Ti extrusions 

seems to produce mechanical properties similar to 

those of cast Ti. New alloys have been targeted that 

could lower the cost of Ti alloys, enabling this class 

of material to enter new markets. Preliminary arc 

melt castings have shown promising microstructures 

and results, and alloy blending practices for the new 

powders are well under way. Further processing development will lead to new low-cost Ti alloy components for heavy vehicle usage. 

References 

1. K. Faller and F. H. Froes. “The Use of Titanium 

in Family Automobiles: Current Trends,” JOM 

53(4), 27–28 (2001) . 

2. F. Froes et al. “Titanium in the Family Automobile: The Cost Challenge,” JOM 56(2), 40–44 

(2004). 

3. J. Rivard et al. “The Thermomechanical Processing of Titanium and Ti-6Al-4V Thin Gage 

Sheet and Plate,” JOM 57(11), 58–61 (2005). 

4. J. Kraft, Summary of Emerging Titanium Cost 

Reduction Technologies: A Study Performed For 

U.S. Department of Energy and Oak Ridge National Laboratory, Subcontract 4000023694, 

Department of Energy, 2004. 

5. B. Hurless et al. “Lowering the Cost of Titanium,” The AMPTIAC Quarterly 6(2), 3–9 

(2002). 

6. P. Blau et al. “Tribological Investigation of Titanium-Based Materials for Brakes,” submitted 

for review. 

7. A. Hartman et al., JOM 50(9), 19 (1998). 

8. H. Friedrich et al., “Titanium in Automotive 

Applications—Nightmare, Vision, or Reality,” 

pp. 3393–3402 in Proceedings, 10th World Conference on Titanium, ed. G. Lutjering, Weinheim, Germany, Wiley-VCH, 2003. 

53
FY 2006 Progress Report High Strength Weight Reduction Materials 

9. U.S. Department of Energy, FY 2005 Progress 

Report for High Strength Weight Reduction Materials, Office of Energy Efficiency and Renewable Energy, DOE, Washington, D.C., 2006. 

54
High Strength Weight Reduction Materials FY 2006 Progress Report 

E. Durability of Titanium Chassis Components 

Principal Investigator: P. J. Blau 

Oak Ridge National Laboratory 

P.O. Box 2008, Oak Ridge, TN 37831-6063 

(865) 574-5377; fax: (865) 574-6918; e-mail:blaupj@ornl.gov 

Technology Development Area Specialist: James J. Eberhardt 

(202)586-9837; fax: (202)587-2476; e-mail: james.eberhardt@ee.doe.gov 

Field Technical Manager: Philip S. Sklad 

(865) 574-5069; fax: (865) 574-6098; e-mail: skladps@ornl.gov 

Contractor: Oak Ridge National Laboratory 

Contract No.: DE-AC05-00OR22725 

Objectives 

• Investigate and develop technology for enhancing the surface durability of titanium (Ti) alloy components, 

which will, in turn, enable their use in lightweight vehicles. 

Approach 

• Design and build a system to simulate road-vibration–induced fretting damage in Ti chassis components such 

as leaf springs. 

• Design and conduct experiments that simulate the kinds of random loads to which springs and fasteners are 

subjected when a vehicle operates. 

• Quantify the fretting damage and identify surface engineering treatments that can be applied to mitigate that 

damage. 

• Determine the effects of load, oscillation amplitude, and joint configuration on the fretting response of 

candidate surface engineering treatments. 

• Provide guidelines to designers of lightweight chassis components regarding engineering solutions to address 

contact damage that can arise when Ti parts are used in wear-critical locations. 

• Establish the relationships between the method of machining Ti parts and their resistance to contact damage. 

Accomplishments 

• Designed and built fixtures on which to conduct fretting tests of flat strips of Ti alloys in a configuration 

representative of leaf springs. 

• Conducted preliminary experiments to develop an understanding of the kinds of damage that can be produced 

on Ti using the new apparatus. 

• Conducted a literature review of fretting-prevention treatments developed for aerospace Ti alloys. The 

information will be used to devise similar approaches for increasing the durability of Ti chassis components. 

55
FY 2006 Progress Report High Strength Weight Reduction Materials 

Future Direction 

• After verifying the characteristics of the testing fixtures and selected protocols, complete baseline studies of 

untreated Ti alloys to provide a basis on which to compare the relative effectiveness of selected frettingprevention methods. 

• Use Oak Ridge National Laboratory (ORNL) cumulative dissipated energy analyses, developed for fatigue 

testing, to assess the relationship between the frictional energy dissipated during fretting contact and the degree 

of damage produced by random-spectrum loading. 

• Evaluate the potential for ORNL-developed oxygen thermal-diffusion treatments in reducing or eliminating 

fretting damage. 

• Prepare Ti surfaces by common industrial machining processes and determine the extent to which the artifacts 

left by those processes can affect the durability of Ti chassis components, both surface treated and untreated. 

Introduction 

Compared with other lightweight alloys, such as 

those based on aluminum and magnesium, Ti alloys have 

higher strength and much greater resistance to the effects 

of elevated temperatures, sustained loads (“creep”), and 

corrosion. While not as low in density as aluminum and 

magnesium alloys, Ti can provide significant 

lightweighting for fuel-efficient automobiles, trucks, and 

trailers while offering durability and corrosion resistance 

as well. 

Recent breakthroughs in primary Ti processing and 

low-cost powder production methods promise to reduce 

material cost. Therefore, the potential exists to expand 

the alloys’ applications beyond aerospace and into 

components for ground transportation. Designers of 

truck and automotive chassis parts such as leaf springs 

have not considered Ti as a viable candidate in the past 

largely because of its cost, but as new processing 

methods bring the price down, it becomes important to 

establish the technology base needed to support the 

broader use of Ti alloys. 

The goal of this new project is to investigate and 

develop technology for enhancing the surface durability 

of Ti components, which will, in turn, expand their 

applications in lightweight vehicles. While there are a 

number of areas that could be addressed, the present 

effort focuses on measuring the resistance of Ti alloys to 

fretting damage, as might be experienced in Ti leaf 

springs and mechanical fasteners. Leveraging past work 

by aerospace engineers, the current approach is to use 

state-of-the-art surface engineering tools to reduce that 

damage and demonstrate their feasibility. The results of 

this work are expected to apply both to the performance 

of Ti-based components in service and to the 

manufacturing of Ti parts. 

Approach 

Fretting is short-amplitude oscillatory motion that 

results in the generation of minute wear particles. The 

cumulative effects of fretting damage include the 

loosening of joints and the initiation of fatigue cracks 

that can eventually lead to catastrophic failures. There 

are a number of possible approaches to fretting damage 

remediation, depending on the kind of mechanical 

configuration to which the materials are exposed. The 

approaches include (1) tightening the clamp load on the 

joint to stop vibration, (2) reducing the load to lower the 

contact pressure on the mating surfaces, (3) lubricating 

the contact to lower the shear forces on the surfaces, and 

(4) substituting more fretting-resistant materials or 

surface treatments. 

Historically, most fretting test apparatuses used in 

fundamental research either use small pads clamped on 

the sides of a flat tensile specimen in a fatigue machine 

or a small sphere of hard material oscillating against a 

flat coupon. Neither of these configurations simulates 

the fretting of flat strips against one another, as might be 

present in leaf springs; therefore, a new test 

configuration was required. 

Once the fixtures were designed and machined, it 

was necessary to characterize the test method itself and 

to determine how best to quantify the fretting resistance 

of selected surface treatments for Ti. The next step was 

to acquire baseline data against which the effectiveness 

of the surface treatments will be judged. Work in 

FY 2006 was therefore focused on preparing the testing 

equipment and conducting preliminary experiments to 

establish a testing protocol. 

56
High Strength Weight Reduction Materials 

Progress 

A flexible test configuration was designed so that 

several possible fretting geometries could be tested using 

slight variations of the same system. The basic fixture 

arrangement is shown in Figure 1. The upper and lower 

components use attachment hardware that is compatible 

with servo-hydraulically actuated mechanical testing 

systems at ORNL. Figure 1 shows a cylindrical upper 

specimen (viewed end-on) and a cantilevered lower 

sheet specimen below it. Repetitive vertical 

displacement of the lower specimen causes fretting to 

occur on the surfaces of both top and bottom specimens. 

The relative amount of fretting depends on the applied 

conditions and whether the mating materials are the 

same or dissimilar. 

Figure 1. This cylinder-on-flat fretting 

configuration was designed to fit existing 

fatigue testing frames. The cylinder specimen 

diameter is 9.53 mm. 

In addition to the cylinder-on-flat contact geometry 

shown in Figure 1, additional contact configurations 

shown in Figure 2 are under consideration. These will 

enable the configuration to simulate either multiple-layer 

springs or pad-on-sheet configurations that can also 

occur in various chassis locations, such as leaf spring 

supports. 

FY 2006 Progress Report 

Figure 2. Three fretting configurations that can 

use the fixture shown in Figure 1. Top: cylinderon-flat sheet. Center: curved sheet-on-flat sheet. 

Bottom: laminar stack of flat sheets. 

Initial studies used both tool steel and Ti as the 

upper cylindrical specimens to fret against Ti-6Al-4V 

sheet specimens. An optical photomicrograph of a sheet 

specimen (Figure 3), subjected to three million cycles 

against a Ti cylinder, shows features that are 

characteristic of oscillatory damage. These include 

abrasion grooves and localized accumulations of 

oxidized debris particles. After the standard testing 

protocol is established, the kinds of damage features will 

be categorized by optical and scanning electron 

microscopy, along with the composition of the debris 

particles. 

57
FY 2006 Progress Report High Strength Weight Reduction Materials 

Figure 3. Optical photomicrograph of the fretting 

surface of a Ti sheet specimen that sustained 

3,000,000 cycles. The sliding direction was vertical 

and the amplitude of sliding was approximately 300 

μm. 

The spectrum of vibrations to which an on-road 

vehicle is subjected depends on its weight and 

suspension design, its vehicle operating profile, and the 

condition of the tires and road surfaces. Fretting damage 

can occur in different ways, depending on the interface 

geometry of the mating parts and the loading conditions. 

A special software program, developed by Don Erdman 

of ORNL, produces a random loading spectrum and 

enables load and displacement data to be obtained. This 

software will be used to simulate road-induced loading 

spectra while collecting data on the cumulative energy 

expended during fretting contact. 

After millions of cycles of contact have been applied 

to the specimens, the resulting surface damage will be 

correlated with the energy used to create it. Once 

baseline data are established for untreated Ti alloys, 

selected surface treatments will be applied to the 

specimens and the results compared with the baseline. 

These surface treatments are planned to include a novel 

diffusion treatment being developed by Jun Qu of 

ORNL. The objective of this aspect of the work is to 

identify the most cost-effective treatments that can be 

used to reduce contact damage on Ti components and 

share that information with manufacturers who are 

interested in developing lightweight chassis components. 

Effects of Surface Grinding on Durability 

The durability of surfaces can be affected not only 

by their surface finish but also by the machining and 

finishing processes used to create them. Machining can 

produce work-hardening of metal surfaces, leave 

residual stresses in the subsurface regions, and introduce 

a degree of directionality (texture) to the microstructure 

that in turn may affect its resistance to contact damage. 

Therefore, a study of the effects of machining damage 

on the fretting and wear resistance of Ti surfaces is 

planned to supplement this effort. 

Future Plans 

During the coming months, the fretting test protocol 

and method of quantifying the extent of surface damage 

will be standardized. Then baseline tests of Ti alloys will 

be conducted. Surface treatment and coating methods 

will be selected and test coupons will be obtained. 

Finally, a study of the effects of grinding conditions on 

the durability of Ti alloy surfaces will be initiated. 

Conclusions 

Application of lightweight Ti alloys for chassis 

components such as leaf springs can be accelerated by 

using what has been learned from prior aerospace 

research and development efforts. Multi-configuration 

fixtures have been prepared, and software to conduct and 

analyze spectrum loading fretting wear tests has been 

obtained. Baseline test protocols are being established, 

and a list of promising candidate surface treatments, 

including one developed at ORNL, is being prepared. 

58
High Strength Weight Reduction Materials FY 2006 Progress Report 

4. ENABLING TECHNOLOGIES 

A. Joining of Advanced Materials by Plastic Deformation 

Principal Investigator: J. L. Routbort (co-workers: D. Singh, Cinta Lorenzo-Martin, and G. Chen) 

Argonne National Laboratory 

9700 S. Cass Avenue, Argonne, IL 60439-4838 

630-252-5065 routbort@anl.gov 

Chief Scientist: James J. Eberhardt 

(202) 586-9837; fax: (202) 587-2476; e-mail: james.eberhardt@ee.doe.gov 

Field Technical Manager: Philip S. Sklad 

(865) 574-5069; fax: (865) 574-6098; e-mail: skladps@ornl.gov 

Contractor: Argonne National Laboratory 

Contract No.: W-31-109-Eng-38 

Objective 

• Join advanced materials such as ceramics, cermets, intermetallics, composites, and biomaterials by plastic deformation and collaborate with industry and universities to produce sensors. 

• Investigate grain rotation that is assumed to occur during grain–boundary sliding (GBS), resulting in deformation bonding. 

Approach 

• Apply a modest compressive load to two pieces of similar or dissimilar materials that have had little surface 

preparation in the temperature region where the materials are known to deform by grain-boundary sliding. 

• Measure the strength of the interface by 4-point bend tests and compare the results with theory. 

• Use electron back-scattered diffraction to measure grain rotation as a function of strain. 

• Prepare for in-situ Advanced Photon Source (APS) experiment to measure grain rotation during deformation. 

Accomplishments 

• Made strong, pore-free joints with various ceramics, cermets, intermetallics, composites, and more recently, 

biomaterials, with and without various interlayers. Fracture occurs away from interface. 

• Published 18 journal papers, obtained patent for plastic joining process, filed patent applications for oxygen 

sensor and joining of advanced materials. 

• Won R&D 100 Award for oxygen sensor. 

Future Direction 

• Join metals and intermetallics to ceramics for solid-oxide fuel cell applications. 

• Complete installation of equipment in beamline to use the Advanced Photon Source to measure grain rotation 

during deformation. We are currently using various routes to produce dense, fine-grained ceramics to deform 

ex-and in-situ. 

59
FY 2006 Progress Report High Strength Weight Reduction Materials 

Foreword 

Progress on this work has not been as rapid as 

anticipated as a result of decreased funding, the 

complexities of designing an in-situ experiment to 

measure grain rotation during deformation of a ceramic at the APS, and repairs on a 30-year old Instron with a high-temperature furnace. However, the 

MTS machine for the APS has been purchased and 

is on site, grips are being designed, and the X-ray 

absorption of a model ceramic has been measured. It 

was determined that we can use a sample of 5-mm 

thickness. Preparations to produce a dense, finegrained NaCl sample continue. These samples will 

be deformed to determine that GBS occurs, and then 

deformed in-situ at the APS while the change of 

grain orientation is measured using diffraction techniques as described in previous annual reports. 

The following discussion is taken from a keynote lecture presented at the conference Mechanical 

Properties in Advanced Materials: Recent Insights, 

Fuenteheridos, Spain, June 7–11, 2006, to be published in the Journal of the European Ceramic Society. It forms the preliminary study for the in-situ 

experiment at the APS scheduled for Fall 2006. 

Introduction 

Study of the high-temperature deformation behavior of polycrystalline ceramics is important from 

several perspectives: (1) high-temperature structural 

applications of the ceramic,

1

 (2) fabrication of the 

ceramics with improved properties,1–3

 and recently, 

(3) joining of polycrystalline ceramics by the plastic 

deformation process.4–6 In general, high-temperature 

deformation of polycrystalline ceramic materials, 

with equiaxed grains, is controlled by a diffusioncontrolled GBS mechanism.

7,8

 It is expected that if 

there is no cavitation and no change in grain shape, 

then the grains must rotate to accommodate the deformation strains.9 

Understanding grain rotation phenomena during 

high-temperature deformation has implications for 

the processing of ceramic materials with enhanced 

mechanical properties, as well as for the joining of 

ceramics. It is possible that minimizing GBS and 

rotation could enhance the creep resistance properties of a material. Recently, it has been demonstrated that various ceramics and intermetallics can 

be joined by high-temperature plastic deformation 

that results in interfaces that are indistinguishable 

from the bulk material.4–6,8,10 This joining technique 

operates by interpenetration of grains at joining surfaces as a result of GBS and grain rotation. There 

are no known experimental studies that have quantified grain rotation in polycrystalline ceramics as a 

function of deformation strains. Thus a hightemperature deformation study has been conducted 

on a model material, strontium titanate (SrTiO3). 

Dense polycrystalline SrTiO3 can be fabricated with 

relative ease and is easily deformed. 

Polycrystalline perovskite materials such as 

BaTiO3 and SrTiO3 are suitable for applications as 

capacitor materials because of their ferroelectric 

behavior.11 Moreover, the Earth’s lower mantle is 

primarily composed of perovskite (Mg, Fe)SiO3 and 

understanding of its plastic deformation behavior is 

important.12 High-temperature deformation behavior 

of BaTiO3 has been extensively studied. Park et al.13 

conducted a high-temperature (1200–1300°C) compressive deformation study on polycrystalline BaTiO3 with fairly large grain sizes ranging from 19– 

52 μm. Their data indicated a stress exponent of 1 

and activation energy of 720 kJ/mol. Moreover, 

Park et al. concluded that the dominant deformation 

mechanism was GBS assisted by lattice cation diffusion. Fine-grained (0.45 μm) BaTiO3 deformed 

superplastically between 1150 and 1250°C with an 

activation energy of 800–1200 kJ/mol.14 

High-temperature deformation literature data for 

SrTiO3 are fairly limited, and most of the studies 

are conducted on single crystals. Using compressive 

tests on SrTiO3 single crystals between 1200 and 

1520°C, Wang et al.12 showed power-law creep behaviour in SrTiO3 for samples compressed in the 

<100> and <110> directions. Creep along the 

<110> direction was much easier than along the 

<100> direction. The stress exponent (n) was ≈3.5 

and activation energy (Q) ranged from 620 kJ/mol 

for the <110> samples and was ≈750 kJ/mol for the 

<100> samples. For the <100> samples, Q was a 

function of stress and/or temperature. 

In this study, we have first investigated the 

high-temperature compressive deformation behavior 

of polycrystalline SrTiO3 at various strain rates in 

the temperature range of 1200–1345°C. Creep 

deformation data and electron microscopy were 

used to ascertain that GBS was the dominant 

deformation mechanism. Subsequently, the electron 

back-scattered diffraction technique was used to 

measure grain rotation as a function of applied 

plastic strain. Measured grain rotations were 

60
High Strength Weight Reduction Materials 

compared with a simulation available in the 

literature.15 

Experimental Details 

The SrTiO3 specimens were fabricated using 

commercially available 98.5%-pure SrTiO3 powder 

(#72117, Alfa Aesar, Ward Hill, MA). The SrTiO3 

powder compacts were formed by uniaxial pressing 

(45 kN force) using a hydraulic laboratory press 

(Carver Laboratory press, Menominee Falls, WI) in 

a cylindrical hardened steel die. The die formed pellets of 2.2 cm diameter and 2.54 cm length. No 

binders were added to the powders prior to pressing. 

The specimens were fired in a box furnace 

(Lindberg Furnaces, Asheville, NC) at 1450°C in air 

for 5 h. The heating rate averaged 3°C/minute, and 

the cooling rate was approximately 4°C/minute. 

The density of the sintered SrTiO3 sample was 

determined by the Archimedes method. Sample 

preparation for microstructural examination involved polishing one of the sample surfaces down to 

a 0.5-μm finísh using alumina slurry. To reveal 

grain boundaries, samples were thermally etched at 

1200°C for 0.5 h in air. Subsequently, microstructure was examined in a Hitachi S-4700-II (Tokyo, 

Japan) field emission scanning electron microscope 

(FE-SEM). Grain size was measured by the standard 

linear intercept method16 using at least 250 intercepts on SEM photomicrographs. Fabricated SrTiO3 

samples were analyzed using inductively coupled 

plasma–atomic emission spectroscopy (ICP-AES) at 

Argonne’s Analytical Chemistry Laboratory to determine the ratio of strontium and titanium. 

Samples for high-temperature mechanical testing were cut, ground, and polished to ≈4×5×7 mm 

parallelepipeds. Constant crosshead-displacementrate experiments were carried out in an Instron machine (Model 1125, Canton, MA) equipped with an 

atmosphere-controlled high-temperature furnace. 

Details of the experimental setup are described 

elsewhere.

5

 The samples were compressed between 

two alumina platens, with platinum foil used as a 

diffusion barrier. Temperatures of 1200–1345oC 

were used, with initial strain rates ranging between 

2×10–6/ s

−1

 and 5×10−5/ s

−1

. Prior to the start of the 

experiment, the chamber was backfilled and flushed 

with argon gas at least three times. All experiments 

were conducted in ≈1 atm of high-purity (99.999%) 

argon atmosphere. At a specific test temperature, for 

each specimen, crosshead speeds were changed 

FY 2006 Progress Report 

(with and without unloading the sample) to obtain 

various nominal strain rates. Strain rates were 

changed only after a steady state (no work hardening) was reached on the load-time trace. 

Microstructural analyses of pre- and postdeformed samples were further conducted using 

SEM and transmission electron microscopy (TEM). 

TEM foils of samples were prepared by slicing thin 

sections using a precision diamond saw. Mechanical 

polishing further thinned samples. Subsequently, 

samples were ion milled using a Gatan (Pleasanton, 

CA) Model 691 precision ion polishing system. The 

microstructural analysis was performed using a 

Philips CM-200 transmission electron microscope 

operating at 200 kV in the Electron Microscopy 

Collaborative Research Center at Argonne National 

Laboratory. 

Quantitative rotations of specific surface grains, 

as a function of applied strain, were measured using 

the electron back-scatter diffraction (EBSD) technique.17 A parallelepiped-shaped SrTiO3 sample, 

with dimensions similar to those of the deformation 

study samples, was used. However, in this case, one 

of the four faces (parallel to the loading direction) 

was polished to a 0.05-μm finísh using alumina polishing solution and then plasma cleaned. A small 

Vicker’s indent was made approximately in the center of the polished face. This was done to provide a 

reference point so that the grains that were followed 

as a function of strain could be easily located. Subsequently, the polished surface was coated with a 

thin layer of carbon. 

EBSD pattern acquisition from the SrTiO3 sample was conducted on a field emission electron microscope (Hitachi S-4500, Tokyo, Japan) equipped 

with an orientation imaging microscopy (TSL, 

Draper, UT) feature at Northwestern University’s 

Electron Probe Instrumentation Center. A special 

sample holder was designed so that the sample orientation in the microscope was always kept the 

same. Sample tilt with respect to the horizontal was 

70°. Diffraction patterns or Kikuchi bands were recorded from ≈20 specific grains. To determine the 

grain rotation, Kikuchi bands were represented by 

spots using the Hough transform.

18 The principle of 

the Hough transform is that the intensity at a point 

in the Hough space represents the intensity along a 

line in the original image. The coordinates in Hough 

space are usually taken as ρ and θ. Here a point (ρ, 

θ) in the Hough transform represents a line at an 

angle θ and at a distance ρ from the center of the 

61
FY 2006 Progress Report 

original pattern. Thus any rotation in the grains will 

rotate the Kikuchi bands and spots. Grain rotation 

was determined by the shift in the position of the 

spots in θ coordinate in the 2-dimensional Hough 

image as a function of the applied strain. If we assume the uncertainty of θ in the Hough transformed 

image is ±1 pixel, the uncertainty in the rotation of 

the Kikuchi line (and the grain) in the original image will be ±1°. This procedure is automated in the 

orientation imaging microscope and was individually applied to all the identified grains. 

Grain rotation measurement, as a function of 

applied strain, was conducted in an interrupted test 

procedure. After ≈20 grains were identified and indexed on the undeformed sample, the sample was 

plastically deformed in steps of 0.02 strains at 

1280°C and at a strain rate of 5×10–6/s. At each step, 

the sample was removed from the Instron and 

placed back in the SEM to collect EBSD patterns 

from the identified grains. This procedure was repeated until the cumulative plastic strain in the sample was 0.08. In addition to tracking the changes in 

the grain orientation with deformation strains, the 

form factors [(4p*grain area)/(grain diameter)2

] of 

the individual grains were tracked with strain. Form 

factor was determined by using standard imaging 

software on an SEM image of the specific grain. 

Results 

The density of the as-fabricated SrTiO3 sample 

was ≈95% of the theoretical density. Examination of 

the microstructure indicated that the grains were 

relatively equiaxed with few entrapped pores. Average grain size was ≈6 μm. 

Major impurities were zirconium, silicon, barium, aluminum, and calcium. Based purely on the 

weight fractions of strontium and titanium, it appears that sintered ceramic is deficient in titanium 

with an Sr/Ti atomic ratio of 1.015. It should be 

noted that zirconium is a major impurity (3.9 wt %). 

A typical stress-strain profile obtained in compression at 1250°C is presented in Figure 1. For the 

most part, at each strain rate, the stress reached a 

steady-state value. Similar responses were obtained 

in deformation tests conducted at other test temperatures. However, at the highest strain rate of 10–4/s, at 

1200°C, there was some evidence of work hardenHigh Strength Weight Reduction Materials 

Stress (MPa) 

10 

8 

6 

4 

2 

0 

2x10-6/s 

5x10-6/s 

1x10-5/s 

0 0.005 0.01 0.015 0.02 0.025 

Deformation 

Figure 1. Stress-deformation profile 

for SrTiO3 obtained during compression testing at 1250°C. 

ing. In such cases, the stress at which plastic deformation was observed was taken as the flow stress. 

Figure 2 represents the strain-rate dependence 

of the flow stress on a log-log basis for the four 

temperatures investigated. Assuming no substantive 

changes occurred in the microstructures during testing, a steady-state phenomenological creep equation 

is used to define the dependence of strain rate on the 

flow stress:19 

ε& = Aσn exp(−Q/RT) (1) 

where ε& is the strain rate, σ is the flow stress, n is 

the stress exponent, A is a constant, Q is the activation energy, R is the gas constant, and T is the absolute test temperature. A stress exponent of n=1 

indicates that diffusion-controlled viscous flow is 

the likely deformation process.1

 Values of n, estimated from best fits of the experimental data to 

Eq. (1), were 0.9 ± 0.1 at 1345oC, 0.9 ± 0.1 at 

1300oC, 1.0 ± 0.1 at 1250°C, and 1.6 ± 0.1 at 

1200oC, as shown in Figure 2. Within the experimental uncertainties in the stress and strain rates, it 

is safe to assume that n = 1 for the three test temperatures from 1250 to 1345°C. 

At a constant strain rate, activation energy, Q, 

for the deformation process can be represented as 

Q = nR ⎜

⎜

⎛

∂

∂

(

ln

1 

σ

)⎟

⎟

⎞

( 2 ) 

⎝ T ⎠ ε&

62
Strain Rate (1/s) 

High Strength Weight Reduction Materials 

0.0001 

10-5 

10-6 

1200°C 

1250°C 

1300°C 

1345°C 

n = 0.9 

n = 1.6 

n = 0.9 

n = 1.0 

0.01 0.1 1 10 100 

Stress (MPa) 

Figure 2. Strain-rate dependency on flow stress for SrTIO3 

samples at various test temperatures. 

Figure 3 shows the Arrhenius plot of flow stress 

as a function of reciprocal absolute temperature at 

four different strain rates tested. Data for tests conducted at temperatures from 1250 to 1345°C were 

considered. From the slopes of the various lines, the 

mean activation energy determined was 628 ± 24 

kJ/mol. It is assumed that the stress exponent, n, is 

constant over the temperature range, and a value of 

n = 1 was used for the calculation. 

Figure 3. Flow stress in SrTiO3 

as a function of test temperature 

at constant strain rates. 

Figure 4 is a typical TEM micrograph of a 

SrTiO3 sample that was deformed at 1350°C with a 

total strain of 5%. The grain boundaries are clean 

with no cavitation observed. No discernible changes 

in the grain shape were observed. In addition, according to the TEM micrograph, no dislocations are 

seen in the deformed sample. However, in the deFY 2006 Progress Report 

Figure 4. TEM micrograph of 

SrTiO3 deformed at 1350°C to e 

= 0.05. Marker is 2 μm. 

formed samples, there is evidence of twins within 

the grains. 

Figures 5a–c show a series of micrographs of a 

particular grain (#24) at strains of 0.02, 0.04, and 

0.06, respectively. These micrographs demonstrate 

that the same grain can be located and tracked in the 

interrupted loading experiment for grain rotation 

measurements. It is clear from these images that 

there is no discernible change in the shape of the 

grain with higher strain. These figures also show the 

diffraction patterns and their respective Hough 

transformed images. The rotation of a grain is detected by the shift in the spots in the Hough space. 

As shown in Figures 5a and 5b, the spot moved by 4 

pixels (from 64 to 68) or the grain rotated by 4° as 

the cumulative strain in the sample was increased 

from 0.02 to 0.04. Subsequently, from 0.04 to 0.06 

strain, the spot shifted by 1 pixel or the corresponding additional grain rotation was ≈1°. 

Figure 6 shows the measurable grain rotations, 

as a function of strain, on 7 grains out of the ≈20 

grains that were followed. It is interesting to see that 

grains rotate in both the positive and negative directions. There is no systematic trend in grain rotation 

with the applied strain. The maximum rotation 

measured was 7°. Furthermore, some grains showed 

immediate rotation at low applied strains and then 

either did not rotate further or rotated back with increasing strains. Form factors measured on some of 

these grains as a function of applied strain indicated 

that there was no significant change in the form factors, within the error of measurement of ±5%, which 

indicates no obvious grain shape change. 

63
FY 2006 Progress Report 

(a) 

(b) 

(c) 

Figure 5. SEM micrograph, diffraction bands, and its 

corresponding Hough transform for a specific grain 

deformed to strains of (a) 2%, (b) 4%, and (c) 6%. 

CUMULATIVE GRAIN ROTATI

ON (°) 

High Strength Weight Reduction Materials 

10 

8 

6 

4 

2 

0 

-2 

-4 

0 0.02 0.04 0.06 0.08 0.1 

GRAIN 7 

GRAIN 11 

GRAIN 14 

GRAIN 15 

GRAIN 20 

GRAIN 23 

GRAIn 24 

DEFORMATION 

Figure 6. Cumulative grain rotations 

as a function of applied strain measured on seven grains in SrTiO3 deformed at 1280°C and 5×10−6

/s. 

Discussion 

Polycrystalline SrTiO3 has been shown to deform in a plastic manner at temperatures >1200°C 

and strain rates as high as 1×10–5/s. This behavior is 

analogous to the high-temperature deformation behavior in other polycrystalline ceramics and composites such as Al2O3 [20], 3Y-TZP (3 mol % Y2O3­

stabilized tetragonal zirconia),2

 and ceramic matrix 

composites.3

 It is generally accepted that the 

mechanism by which deformation or plastic flow 

occurs in these materials is by diffusion-assisted 

GBS.2,3

 Conditions conducive to deformation by 

plastic flow are reasonably high test temperatures 

and fine and equiaxed grain structure to accommodate the displacements of individual grains.7 

The steady-state creep equation, as represented 

in Eq. (1), adequately defines the deformation behavior of SrTiO3 for temperatures >1200°C and 

strain rates ranging from 2×10–6/s–1×10–5/s. A stress 

exponent of n≈1 is indicative of a diffusioncontrolled deformation process.7

 Further, the absence of any deformation-induced dislocation structures in the TEM analysis (as shown in Figure 4) 

indicates that SrTiO3 deforms by GBS. Park et al. 

also observed no change in grain shapes in their 

study of a relatively large-grain (19–50 μm) BaTiO3 

ceramic and concluded that high-temperature deformation was controlled by GBS.13 The stress exponent in their study was reported to be ≈1. The 

higher stress exponent value (n=1.6) for tests on 

SrTiO3 conducted at 1200°C is not clear. This observation could be due to the onset of microfracture 

or a change in the deformation mode. 

64
High Strength Weight Reduction Materials FY 2006 Progress Report 

Insight into the possible deformation mechanism(s) can be obtained from the activation energy 

measurements. Activation energy for deformation of 

SrTiO3 ceramic was 628 ± 42 kJ/mol, as determined 

from the Arrhenius plot of Figure 3. This value is 

quite similar to an activation energy of 620kJ/mol 

reported by Wang et al.12 on the creep deformation 

of a SrTiO3 single crystal deformed along {100} 

<010> orientation. The stress exponent for singlecrystal SrTiO3 was 3.5, and the deformation mechanism was controlled by diffusion-assisteddislocation climb and was independent of temperature. Furthermore, Wang et al.12 point out that this 

activation energy is similar to that of the diffusion 

of Ti+4 ions, which is the slowest and ratecontrolling species in the titanate perovskites.21,22 

For layered perovskites, such as superconductors, it has been shown that creep is controlled by 

diffusion of A site cations.23,24 For SrTiO3, the Asite cation is an Sr ion. The activation energy of Sr 

diffusion in polycrystalline SrTiO3 is reported to be 

293 kJ/mol,25 which is significantly lower than the 

activation energy determined from the present 

study. It is quite possible that, in polycrystalline 

SrTiO3, diffusion of the B-site cation or titanium ion 

is the deformation controlling species because of its 

+4 valence state and octahedral coordination with 

oxygen ions, as suggested by theoretical defect studies and experimental measurements for Ti

+4 diffusion in titanates.21,22

 A similar possibility that 

diffusion of B-site cation or titanium ion is the deformation controlling specie has been suggested by 

Park et al. in their work on polycrystalline BaTiO3.

13 

Furthermore, the role of zirconium ion impurity (≈4 

wt %) in the deformation and hence the measured 

activation energy was not explored. 

The activation energy determined for polycrystalline SrTiO3 is on the lower end compared with 

the activation energies of other complex-structured 

perovskite-type ceramics, as shown in Table 1. 

Normalized activation energy [Q/RTm, (R=gas constant and Tm=melting point)] allows one to compare 

the activation energy for various other material systems. Moreover, for these complex oxides, it is 

known that n≈1 and the deformation mechanism is 

GBS. It is clear from Table 1 that there is a wide 

distribution for the normalized activation energies 

for the various perovskites. 

Table 1. Activation energies for deformation of various 

oxide ceramics 

Material Q Q/RTm Reference 

(kJ/mol) 

BaTiO3 720 46 Park et al. 

(13) 

CaTiO3 840 45 Yamada (33) 

SrZrO3 710 29 Nemeth (34) 

SrTiO3 

single 750 39 Wang et al. 

crystal 620 32 (12) 

<100> 

<110> 

SrTiO3 628 33 This work

There are numerous examples of observance of 

grain rotation during deformation, grain growth26 or 

texture evolution in metallic systems.27 It is generally accepted that during plastic deformation, grain 

rotation is a consequence of a net torque on the 

grain.15 However, in concert with grain rotation, an 

accommodating process needs to occur so that no 

voids or cavities are created. This accommodation is 

accomplished by GBS that, in the case of SrTiO3, is 

assisted by cation diffusion. 

Evidence of grain rotation in SrTiO3 ceramics 

as a function of applied strain is clear from Figures 

5 and 6. From SEM and TEM analysis of the deformed samples there was no evidence of cavitation. 

In addition, the form factors or the grain shapes of 

the specific grains for which rotation was measured 

did not change within the accuracy of the experimental measurements. Thus to accommodate the 

plastic strains, the grains must rotate.9

 Ashby and 

Verrall9

 have shown that plastic strains can be accommodated by diffusion-assisted GBS that involves grain translation and rotation and little or no 

change in grain shape or size. Based on the AshbyVerrall model, large strains are accommodated by 

neighbor-switching events, and the degree of grain 

rotation is dependent on the surrounding grains.9 

This model is fundamentally different from Nabarro-Herring28 and Coble29 creep models that predict grain shape changes with no grain rotations. 

Gifkin3 ̊ pointed out that strain rate predictions 

from the Ashby-Verrall model were on the lower 

end of the experimental measurements and that the 

grain size dependence did not agree with the experimental observations. Further, a simple 2­

dimensional Ashby-Verrall model does not account 

for the changes in the surface area. Gifkin3 ̊

 showed 

65
FY 2006 Progress Report 

in his model that changes in the surface area can be 

accounted for by considering the grains emerging 

from subsurface or surface grains disappearing into 

the bulk. This effect was first observed and reported 

by Rachinger.31 Gifkin’s model3 ̊

 also predicts grain 

rotation. Evidence of grain disappearing into the 

bulk was also observed in the present study. Thus it 

appears that in addition to diffusion-assisted GBS, 

there is movement of grains from subsurface to interior and vice versa to account for the surface area 

changes as predicted by Gifkin.3 ̊

Recently, there have been several meso-scale 

simulation efforts to quantify grain rotation during 

plastic deformation. Notably, Ding et al.32 have 

simulated grain rotation for a model material with 

5000 grains. They assumed equiaxed grains with a 

log normal grain size distribution and varying diffusive fluxes. However, no grain sliding resistance 

was considered. Grain-boundary diffusion or Coble 

creep accommodated deformation. They found that 

at 50% plastic strain, grain rotations ranged from 

−10° to +10°, with almost 40% of the grains showing no rotation. Another 40% of the grains showed 

rotations between −5° and +5°. These simulations 

are qualitatively similar to the experimental observations made on grain rotations as a function of 

plastic strain in SrTiO3 in this study. It is recognized 

that the number of grains followed for rotation in 

this study are too small to provide a statistically 

meaningful comparison. 

Grain rotation measurements were conducted in 

an ex-situ manner. Reloading of the specimen at 

each strain level is expected to create a somewhat 

different stress state at the microstructure level. In 

addition, the rotations measured are only for the 

surface grains. It is expected that grains in the sample interior will be exposed to different stresses as 

well as different diffusion conditions. In this regard, 

an in-situ experiment will be definitive where the 

grains in the bulk can be tracked for rotation. 

Conclusions 

High-temperature compressive deformation behavior of a polycrystalline strontium titanate 

(SrTiO3) ceramic has been investigated between 

1200 and 1345°C at strain rates ranging from 5×10– 

6

/s to 5×10–5/s. Stress exponents of ≈1 indicated a 

viscous diffusion-controlled deformation mechanism with an activation energy of ≈628 kJ/mol. 

GBS was identified as the principal deformation 

High Strength Weight Reduction Materials 

mechanism. Comparison of activation energy with 

literature data on single-crystal SrTiO3 suggests that 

diffusion of titanium ions is rate controlling. The 

EBSD technique was used to determine grain orientation as a function of applied strain in SrTiO3. 

Grain rotations were both positive and negative, and 

rotations as large as 7° were measured. The magnitudes of grain rotation angles were qualitatively 

similar to those predicted from simulations. 

References 

1. W. R. Cannon and T. G. Langdon, “Creep of 

Ceramics—1. Mechanical Characteristics,” J. 

Mater. Sci. 18 (1), 1–50 (1983). 

2. F. Wakai, S. Sakaguchi, and Y. Matsuno, “Superplasticity of Yttria-Stabilized Tetragonal Zirconia,” Adv. Ceram. Mater. 1, 259 (1986). 

3. A. H. Chokshi, A. K. Mukherjee, and T. G. 

Langdon, “Superplasticity in Advanced Materials,” Mater. Sci. Eng. R10 237 (1993). 

4. J. Ye and A. Domininquez-Rodriguez, “Joining 

of Y-TZP Parts,” Scripta Metall. Mater. 33, 

441–6 (1995). 

5. F. Gutierrez-Mora, A. Dominguez-Rodriguez, J. 

L. Routbort, R. Chaim, and F. Guiberteau, 

“Joining of Yttria-tetragonal Stabilized Zirconia 

Polycrystals using Nanocrystals,” Scripta Mater. 41, 455–60 (1999). 

6. K. C. Goretta, F. Gutierrez-Mora, J. J. Picciolo, 

and J. L. Routbort, “Joining Alumina/Zirconia 

Ceramics,” Mater. Sci. & Eng. A341, 158–162 

(2003). 

7. T. G. Langdon, “The Characteristics of Superplastic-Like Flow in Ceramics,” in Plastic Deformation of Ceramics, eds. R. C. Bradt, C. A. 

Brookes, and J. L. Routbort, Plenum Press, New 

York, pp. 251–268 (1995). 

8. F. Gutierrez-Mora, K. C. Goretta, S. Majumdar, 

J. L. Routbort, M. Grimdisch, and A. Domínguez-Rodriguez, “Influence of Internal Residual 

Stresses in Superplastic Joining of Airconia 

Toughened Alumina,” Acta Mater. 50, 3475– 

3486 (2002). 

9. M. F. Ashby and R. A. Verrall, “DiffusionAccommodated Flow and Superplasticity,” Acta 

Metall. 21, 149–163 (1973). 

10. D. Singh, F. Gutierrez-Mora, N. Chen, K. C. 

Goretta, and J. L. Routbort, “Joining of Advanced Materials by Plastic Deformation,” Ceram. Eng. Sci. 25(3), 113–119 (2004). 

66
High Strength Weight Reduction Materials 

11. F. Jona and G. Shirane, in Ferroelectric Crystals, Dover, New York, 1993 

12. Z. Wang, S. Karato, and K. Fujino, “High Temperature Creep of Single Crystal Strontium Titanate (SrTiO3): A Contribution to Creep 

Systematics in Perovskites,” Phys. Earth & 

Planet. Inter. 79, 299–312 (1993). 

13. E. T. Park, P. Nash, J. Wolfenstine, K. C. Goretta, and J. L. Routbort, “High-temperature 

Creep of Polycrystalline BaTiO3,” J. Mater. 

Res. 14(2), 523–528 (1999). 

14. C. Carry and A. Mocellin, J. Am. Ceram. Soc. 

69, C215 (1986). 

15. D. Moldovan, D. Wolf, and S. R. Phillpot, 

“Theory of Diffusion-Accommodated Grain Rotation in Columnar Polycrystalline Microstructures,” Acta Mater. 49, 3521–3532 (2001). 

16. E. E. Underwood, in Quantitative Stereology, 

Addison-Wesley, Reading, MA (1970). 

17. F. J. Humphreys, “Review: Grain and Subgrain 

Characterization by Electron Backscatter Diffraction,” J. Mat. Sci. 36, 3833–3854 (2001). 

18. X. Tao and A. Eades, “Errors. Artifacts, and Improvements in EBSD Processing and Mapping,” 

Microsc. Microanal. 11, 79–87 (2005). 

19. T. Bretheau, J. Castaing, J. Rabier, and P. 

Veyssiere, “Dislocation Motion and HighTemperature Plasticity of Binary and Ternary 

Oxides,” Adv. Phys. 28, 835 (1979). 

20. L. A. Xue and I. W. Chen, “Deformation and 

Grain Growth of Low-Temperature-Sintered 

High-Purity Alumina,” J. Am. Ceram. Soc. 72, 

3518 (1990). 

21. W. H. Rhodes and W. D. Kingery, “Dislocation 

Dependence of Cationic Diffusion in SrTiO3,” J. 

Am. Ceram. Soc. 49(10), 521–526 (1966). 

22. K. Wright, Rheology of Perovskites and Its Implications for Mantle Dynamics, Ph.D. Thesis, 

Department of Geological Science, University 

College, London (1991). 

23. A. Dominguez-Rodriguez, J. Castaing, K. C. 

Goretta, and J. L. Routbort, in Point Defects and 

Related Properties of Ceramics, eds. T. O. Mason and J. L. Routbort, p. 139, American Ceramic Society, Westerville, OH, (1991). 

24. M. Jimenez-Melendo, A. R. De Arellano-Lopez, 

A. Dominguez-Rodriguez, K. C. Goretta, and J. 

L. Routbort, “Diffusion-controlled Plastic Deformation of YBa2Cu3Ox,” Acta Metall. Mater. 

43, 2429 (1995). 

FY 2006 Progress Report 

25. P. Turlier, P. Bussiere, and M. Prettre, Compt 

Rend. Acad. Sci., 250, p. 1649, Paris (1960). 

26. P. G. Shewmon, in Recrystallization, Grain 

Growth and Textures, ASM, Metals Park, OH, 

1966. 

27. T, Yamaskai, Y. Demizu, and Y. Ogino, Mater. 

Sci. Forum 461, 204–206 (1996). 

28. C. Herring, J. Appl. Phys. 21, 437 (1950). 

29. R. L. Coble, J. Appl. Phys. 34, 1679 (1963). 

30. R. C. Gifkin, “Grain Rearrangements during 

Superplastic Deformation,” J. Mat. Sci. 13, 

1926–1936 (1978). 

31. W. A. Rachinger, J. Inst. Metals 81, 33 (1952– 

53). 

32. R. Ding, D. Moldovan, V. Yamakov, D. Wolf, 

and S. R. Phillpot, “Effects of Microstructural 

Inhomogeneity on Dynamic Grain Growth during Large-Strain Grain Boundary DiffusionAssisted Plastic Deformation,” Modelling Simul. 

Mater. Sci. Eng. 13, 1129–1151 (2005). 

67
FY 2006 Progress Report High Strength Weight Reduction Materials 

68
High Strength Weight Reduction Materials FY 2006 Progress Report 

B. Attachment Techniques for Heavy Truck Composite Chassis Members 

Principal Investigator: Barbara Frame 

Oak Ridge National Laboratory 

P.O. Box 2008, Oak Ridge, TN 37831-6053 

(865) 576-1892; fax: (865) 574-8257; e-mail: framebj@ornl.gov 

Chief Scientist: James J. Eberhardt 

(202) 586-9837; fax: (202) 587-2476; e-mail: james.eberhardt@ee.doe.gov 

Field Technical Manager: Philip S. Sklad 

(865) 574-5069; fax:(865) 574-6098; e-mail: skladps@ornl.gov 

Contractors: Oak Ridge National Laboratory, Pacific Northwest National Laboratory 

Contract Nos.: DE-AC05-00OR22725, DE-AC06-76RL01830 

Objectives 

• Overcome the technical issues associated with joining composite materials in heavy vehicles by developing 

technically robust and economically attractive joining techniques. 

• Develop and validate one or more joint designs for a composite structural member attached to a metal member 

that satisfy the truck chassis structural requirements both economically and reliably. 

• Solicit input from truck original equipment manufacturers (OEMs) and suppliers on the technical hurdles and 

needs associated with joining structural composite members in heavy vehicles. Use this information to guide 

the joint design and development activities. 

• Publish information on the design, modeling, and testing methodologies that are developed to support the incorporation of composite materials into other chassis components. 

Approach 

• Collaborate with the National Composites Center (NCC) and its OEM partners to identify and address technical 

needs related to the manufacturing, joining, and implementation of a composite chassis component. 

• Design attachment components and configurations in close coordination with the composite structural component development. 

• Use modeling techniques to predict the performance of various joint designs, taking into account damage 

mechanisms and fatigue/life requirements. 

• Characterize various composite materials and mechanical joint configurations through mechanical testing, considering variables such as hole size, hole fabrication method, bolt preload, inserts, combined loading, vibration, 

fatigue, and durability. 

• Validate joint design for the composite structural member through shaker and track testing. 

Accomplishments 

• Conducted constant amplitude and block loading fatigue of candidate three-dimensional (3-D) glass fiber reinforced composites. 

• Conducted double lap shear tensile and fatigue tests to investigate changes in bolt preload resulting from 

through-thickness creep of the composite and damage progression. 

69
FY 2006 Progress Report High Strength Weight Reduction Materials 

• Developed data acquisition software to calculate and monitor energy expended per cycle during tension-tension 

fatigue tests and initiated evaluation with commercial-grade carbon fiber composite. 

• Conducted environmental tension-tension fatigue tests of bolted 3-D glass fiber reinforced composite/steel lap 

shear joints under conditions of water and salt water exposure. 

• Conducted finite element analyses (FEAs) of 3-D glass fiber reinforced composite/steel bolted double lap 

joints. 

Future Direction 

• Continue bolt bearing and fatigue tests using thicker (1⁄4 – 1⁄2 in.) fiberglass composites to investigate design 

modifications, adhesive bonding, and acceptable hole clearance levels to minimize damage progression in the 

composite and improve the fatigue life of a composite-to-steel joint. 

• Continue environmental fatigue tests with 3-D glass fiber reinforced composite and incorporate vehicle cleaning fluid and elevated temperature into the environment test matrix. 

• Continue tension-tension fatigue tests with developed data acquisition software and commercial-grade carbon 

fiber composite to evaluate capabilities to estimate fatigue life using an energy dissipation approach. 

• Continue FEA modeling with input from the composite and joint tests to further optimize and predict the performance of the composite/steel joint. 

Introduction 

In May 2003, researchers at Oak Ridge National 

Laboratory (ORNL) and Pacific Northwest National 

Laboratory (PNNL) began collaboration on a 4-year 

research effort focused on developing technically 

robust and economically attractive joining techniques to overcome the technical issues associated 

with joining lightweight materials in heavy vehicles. 

This project has been conducted in close partnership with the NCC and two truck industry OEMs. 

These partners have contributed significantly toward 

many previous accomplishments in the project by 

providing guidance regarding the service environment, manufacturing processes, and end user expectations. Some of the testing work was even done in 

an OEM partner’s plant, so that it could simulate 

real manufacturing conditions as closely as possible. 

These partners were working on a cost-shared, 

partially DOE-funded project that serendipitously 

served as a “focal project” for the team. Major reductions in the HSWR program necessitated a major 

reduction in the DOE funding for the industry project, and the research team experienced a corresponding decrease in interaction with its industry 

partners. Consequently, this year the team has addressed joining issues of general interest, with increased attention to problems that may arise 

throughout the truck chassis, but with decreased industry feedback regarding the practical and specific 

implementations of those solutions. 

Composite Material Testing 

Variable Amplitude Fatigue 

Components in a truck chassis undergo complex 

random variable fatigue loading during their lifetime. Previous studies reported in the literature indicate that standard methods employed for variable 

amplitude fatigue life prediction can overestimate 

the actual experimentally measured life. The inaccuracy is due in part to the different failure mechanisms and property degradation in fiber-reinforced 

composite materials. 

For example, the Palmgren-Miner linear relationship assumes that if a block of loading cycles at 

a given stress amplitude (n1) is followed by a subsequent block of cycles at different stress amplitude 

(n2), that overall damage accumulation is determined 

as a sum of the ratios of the loading cycles (n) to the 

failure cycles (N) at each stress level. Failure is predicted when the sum of the ratios equals 1: 

n1/N1 + n2/N2 + ... = 1 

70
400 

High Strength Weight Reduction Materials FY 2006 Progress Report 

3D glass/epoxy - 30% to 50% 3D glass/epoxy - 50% to 30% 

3D glass/polyurethane - 30% to 50% 

The Palmgren-Miner linear relationship assumes 

3D glass/polyurethane - 50% to 30% that fatigue cycles at a specific load level always 

result in the same damage regardless of the loading 3.50 

history. The literature suggests that the model yields 3.00 

acceptable results for metals and is used for life pre- 2.50 

Miner's rule 

prediction 

dictions in many practical applications for metallic 2.00 

1.50 

characterize the behavior of 3-D glass fiber– 1.00 

reinforced epoxy and 3-D glass fiber–reinforced 0.50 n2/N2 

structures. 

Block loading fatigue tests were conducted to 

polyurethane composite. A comparison of measured 

fatigue life with the Palmgren-Miner’s prediction 

was made and some deviation was evident. 

Quasi-static tensile tests indicated that the 

3-D glass fiber–reinforced polyurethane specimens 

had an average tensile strength of 378.6 MPa and 

modulus of 24.0 GPa. The 3-D glass fiber– 

reinforced epoxy specimens had a slightly higher 

average tensile strength of 421.9 MPa and a 

modulus of 23.8 GPa. 

Figure 1 is a plot of constant amplitude tensiontension fatigue test results for the 3-D glass fiber 

reinforced epoxy and 3-D glass fiber–reinforced 

polyurethane composite. Although the number of 

data points is limited, the plots in Figure 1 indicate 

that the constant amplitude fatigue results of the two 

materials are similar. 

0.00 

0.00 0.25 0.50 0.75 1.00 

n1/N1 

Figure 2. Block loading fatigue results for 3D glass-fiber– 

reinforced epoxy and 3D glass-fiber–reinforced polyurethane matrix specimens. The two-block maximum fatigue 

loads as a percentage of ultimate tensile strength are 50% 

and 30%. 

glass/polyurethane composite data suggest that the 

two-block amplitude load sequence is more damaging for these specimens than would be predicted by 

the Palmgren-Miner model. The 3-D glass fiber– 

reinforced epoxy matrix specimens showed less deterioration of predicted fatigue life from the twoblock amplitude load sequence, and data points from 

the high-to-low sequences exceed predicted life significantly. 

For both epoxy and polyurethane matrix com350 posites, the lower stress level applied prior to the

1.E+02 1.E+03 1.E+04 1.E+05 1.E+06 1.E+07 

Cycles to failure, Nf 

Polyurethane matrix 

Epoxy matrix 

Run out 

Maxim um Stress Level (MPa) 

high stress level was more damaging than when the 

higher stress level was applied first. 

300 

250 

It is hypothesized that the difference in perform­ 200 ance between the polyurethane and epoxy composite 150 specimens with block loading fatigue may be attrib100 utable to the properties of the polyurethane and epoxy matrix resins, the quality of the fiber-matrix 50 

interface, and specimen geometry. The level of sta­ 0 tistical confidence in this study is also shaped by the 

small number of replicates available for each testing 

condition. Figure 1. Constant amplitude fatigue test results for 3D These tests demonstrated that the 3-D glass fiber–reinforced composite specimens can become 

visibly whitened and opaque with evidence of fatigue damage (attributed to microcracks, matrix 

crazing, and failure of the fiber-resin interface) 

glass-fiber–reinforced epoxy and polyurethane matrix 

composite specimens. 

Figures 2 shows block loading fatigue results for 

the 3-D glass fiber–reinforced polyurethane and 3-D 

glass fiber–reinforced epoxy composites. The two- but still retain sufficient load-bearing capability to 

block maximum fatigue loads as a percentage of ul­ continue cycling for some period of time before 

timate tensile strength are 50% and 30%. The 3-D eventual failure. That is, the presence of some visible damage does not signal that catastrophic failure 

71 
FY 2006 Progress Report 

is imminent. It is speculated that the 

3-D reinforcement may be beneficial in this regard 

by helping keep the sample together (intact). 

This research team has been working closely 

with Alpha Star to integrate a damage model into 

Genoa that will take into account the deviation from 

the Palmer-Miner prediction. Additionally, the progressive failure analysis in Genoa may assist in developing an understanding of the damage 

development and evolution which may result in the 

different behavior based on the loading sequence. 

The ultimate goal would be to have a predictive tool 

that could be used for a more realistic random amplitude fatigue spectrum. This would be helpful to reduce the high cost of material characterization 

associated with fatigue testing, but this research is 

beyond the scope of this program. 

Fatigue Energy Dissipation Study 

It is hypothesized that a relatively simple relationship may exist between fatigue amplitude and 

damage accumulation leading up to catastrophic 

failure. Such a relationship could be better used to 

predict structural survival in composites than traditional methods such as S-N curves or the PalmgrenMiner model adopted from high cycle fatigue of 

metals. An exploratory study was therefore initiated 

with a commercially available carbon fiber– 

reinforced cross-ply laminate to investigate the fundamental relationships between energy dissipation 

and damage accumulation, and their utility in making fatigue life predictions. 

A typical load displacement record for an individual fatigue cycle consists of a loading curve and 

unloading curve as depicted in Figure 3. In the case 

of a viscoelastic composite under load control, the 

associated displacement may lag the applied load, 

resulting in a hysterisis loop that will remain constant within the elastic range of the material. However, any damage accumulation should be evident by 

observing changes in specimen stiffness (increased 

displacement) under the constant amplitude loading, 

which will change the area contained within the hysteresis loop. 

The proposed approach assumes that the cumulative damage energy from each cycle is simply the 

work or the area under the load-displacement curve 

for each fatigue cycle summed over the life of the 

High Strength Weight Reduction Materials 

Figure 3. Schematic of load-displacement hysteresis. 

structure. At failure, the total accumulative energy 

and the number of cycles may be easily measured 

from a fatigue test. This total energy would represent the energy capacity of the material in question. 

With this information, it is reasonable to assume that 

the number of cycles to failure may be predicted for 

different fatigue amplitudes and loading scenarios, 

based on the cumulative damage energy results obtained from earlier tests. The key to this approach is 

to determine the energy component responsible for 

damage. 

Although this approach may appear simplistic, it 

requires determining energy levels for each cycle of 

a fatigue test, which may last for millions of cycles. 

This poses a data acquisition and computational 

challenge. The second challenge is determining the 

precise energy component(s) that correlate with the 

damage progression in the composite. 

For example, the total potential energy for a single cycle will include the (1) elastic strain energy, 

(2) energy dissipated due to heating, (3) any plasticlike permanent deformation, and (4) dissipative energy associated with damage. It is the latter of these 

energy terms which must be discerned from each 

load-displacement curve associated with each fatigue cycle. If this can be accomplished, it is reasonable to assume that it would be possible to establish 

an energy capacity for a given material by accumulating all of the dissipative damage energy for a particular material over a series of fatigue cycle 

loadings leading to failure. 

Fatigue tests are being conducted under several 

constant amplitude load ranges with the areas under 

72 
High Strength Weight Reduction Materials 

the loading and unloading curves being recorded for 

every fatigue cycle in the test up until specimen failure. These data can be used to determine the area 

enclosed in the hysterisis loop and other associated 

information to establish any correlation between 

these energy levels and the required number of cycles to failure. Strongly correlated data would suggest a relation between energy accumulation and the 

fatigue life of the material and provide phenomenological evidence to explore any correlation at a more 

fundamental, rigorous level. 

Static and Fatigue Double-Lap, ShearJoint Testing 

A testing methodology has been developed to 

study the effects of bolt preload with composite to 

steel joints. The double-lap, shear-bolted composite/steel assembly in combination with an interface 

load cell to monitor the bolt-clamping force during 

testing is being used to evaluate the effect of 

changes in the torque level on static strength and 

fatigue life. Specimen size is 76.2 mm wide ×

177.8 mm long × 2.5 mm thick. Holes (12.7 mm 

diam) are drilled using a Forstner bit, and assembly 

was with 12.7-mm (1⁄2-in.) grade-8 bolts. The torque 

is applied by hand with a torque wrench adjusted to 

the proper setting. 

Bolt Preload and Quasi-Static Strength 

Quasi-static tensile tests conducted with 2.54– 

3.17-mm thick fiberglass-reinforced composites 

show that the tensile load-bearing capability of the 

composite/steel assembly increases as bolt preload 

increases. Failure modes shift from bearing type 

failures at the lower preloads, typified by local composite crushing at the hole perimeter, to composite 

material tensile failures initiating at the hole. This is 

attributed to the higher clamping forces enabling a 

more effective transfer of tensile load between material adjacent to the hole perimeter and the surrounding (bulk) composite. 

Materials evaluated to date include 3-D glass fiber–reinforced epoxy and 3-D glass fiber–reinforced 

vinyl ester composites molded by the NCC, and a 

commercial pultruded fiberglass Extren® sheet 

product. Figure 4 shows the maximum joint break 

load of specimens prepared with a specified composite material as a function of bolt torque. 

FY 2006 Progress Report 

Break Load (kN) 40 

35 

30 

25 

20 

15 

10 

5 

0 

3D glass/epoxy matrix 

3D glass/vinyl ester matrix 

Extren (pultruded fiberglass) 

Finger-tight 41 122 203 

Bolt Torque (Nm) 

Figure 4. Tensile break loads of bolted double lap shear 

composite/steel assemblies as a function of bolt pre-load. 

Through-thickness Creep and Loss of 

Bolt PreLoad 

An important component of a bolted composite/steel joint design is the bolt preload at assembly, 

which is selected to achieve the desired clamp load 

for proper function of the joint throughout the service life of the assembly. A minimum clamping 

force is required to overcome vibration-induced 

loosening, joint separation, slippage, fatigue, leakage and other similar failures. The maximum clamp 

force should be below a threshold value that would 

cause bolt yielding, joint crushing, stress cracking, 

premature fatigue failure, or other related service 

failures. 

The low through-thickness properties typical of 

most composite architectures suggest that bolted 

composite assemblies will experience a loss of preload over time due to the relaxation or creep of the 

softer composite material. 

Tension-tension fatigue joint tests were conducted with 3-D glass fiber–reinforced epoxy specimens provided by the NCC. For these tests, the 

torque level was maintained at 41 Nm (30 ft-lb). 

This torque nominally produces a bolt clamping 

force of 15–17 kN as measured directly by the interface bolt load cell. The maximum fatigue load was 

established as a percentage of the ultimate tensile 

break load of comparably prepared specimens. The 

fatigue tests were run with a stress cycle R factor of 

0.1 and a frequency of 5 Hz. 

Figure 5 is a plot of the representative loss of 

bolt preload data as a function of the fatigue cycle 

number, and measured for fatigue levels of 70%, 

73
FY 2006 Progress Report High Strength Weight Reduction Materials 

tive of the projected fiberglass chassis component is 104 highly desirable for these investigations. 

102 

Steel baseline Environmental Fatigue Tests

Normalized Preload (%) 100 

98.2

98 3D glass composite 

96 95.1 

94.7 

70% UTS 94 94.6 

80% UTS Grip failure 

90% UTS 92 Steel 

90 

1.E+00 1.E+01 1.E+02 1.E+03 1.E+04 1.E+05 1.E+06 1.E+07 1.E+08 

Cycles, Nf 

Figure 5. Normalized loss of pre-load for 3D glass-fiber– 

Environmental fatigue testing is being conducted to evaluate the durability of bolted composite/steel joints with exposures to water, 

miscellaneous chemical fluids, and temperature extremes likely to be encountered during the long-term 

operation of a truck. Testing is conducted with 

bolted composite/steel lap shear specimens and the 

composite material is the 3-D glass fiber–reinforced 

epoxy composite provided by the NCC. Specimen 

size is 25.4-mm wide × 101.6-mm long × 2.5-mm reinforced epoxy composite and steel baseline specimens 

thick. Holes (6.4-mm diameter) are drilled using a during joint tension-tension fatigue test. Forstner bit and assembly is with 6.4-mm (1/4-in.) 

grade 8 bolts. The torque is 20 Nm applied by hand 

with a torque wrench adjusted to the proper setting. 

Environmental fatigue tests to date have been 

conducted at room temperature in ambient air, distilled water, and 5 wt % salt water environments. 

The maximum fatigue loads are established as a percentage of the ultimate tensile load of comparably 

assembled specimens tested in air and at room temperature. Fatigue tests were run with a stress cycle R 

factor of 0.1 and a frequency of 5 Hz. 

Figure 6 is a plot of the environmental fatigue 

test results to date. The data suggests that water immersion does not significantly reduce fatigue life of 

the 3-D glass fiber–reinforced epoxy composite. 

Specimens immersed in water achieved run-out with 

over 2 million cycles at a fatigue level of 60% to 

70% of the ultimate failure load. Specimens tested in 

5 wt % salt water exhibited a reduced fatigue life 

compared to the specimens tested in air and distilled 

water. 

Comparable 3-D glass fiber–reinforced epoxy 

composite coupons (25.4-mm wide × 101.6-mm 

long × 2.5-mm thick with a 6.4-mm diameter hole) 

were immersed in distilled water, 5 wt % salt water, 

and a commercial car wash solution to measure 

weight change with time immersion. Results are 

summarized in Figure 7. Weight increases of less 

than 0.20% were measured with coupons immersed 

in distilled water and salt water immersions after 

27 days and with coupons immersed in a car wash 

solution for 21 days. This suggests that there is not 

appreciable “wicking” of fluid into the composite 

80% and 90% of the ultimate failure load. Data from 

a steel baseline specimen in which a steel plate was 

substituted for the composite are included for comparison. The data show that the loss of the preload of 

the bolted steel/steel assembly during tensiontension fatigue is considerably less than that of the 

composite/steel assemblies. The higher loss of bolt 

preload measured with the composite assemblies is 

attributed to through-thickness creep of the composite. 

Fatigue joint testing with nominal 2.5-mm 

(0.10-in.) 3-D glass fiber–reinforced epoxy specimens provided by the NCC shows that the double 

lap shear composite/steel assemblies withstand reasonably high fatigue load levels without failure, provided that the initial bolt clamping force is at least 

15–17 kN. For instance, one specimen achieved runout with over 2.6 million cycles at a fatigue level of 

95% of the ultimate failure load. 

Other failures occurred at over 1 million cycles 

at a fatigue level of 90% of the ultimate failure load, 

but it did not initiate at the hole. These failures occurred either in the bulk composite or at the grip. 

This suggests that the fatigue life of the bolted composite/steel assembly is currently on par with that of 

the 2.54-mm fiberglass composite, provided that 

adequate clamping force is maintained for the duration of the test (i.e., no loss of preload occurs due to 

through-thickness creep of the composit. 

Recommendations are that further composite/steel assembly fatigue investigations be conducted using thicker (and more robust) fiberglass 

specimens. A specimen thickness that is representa74 
High Strength Weight Reduction Materials FY 2006 Progress Report 

40 

50 

60 

70 

80 

90 

100 

1.E+02 1.E+03 1.E+04 1.E+05 1.E+06 1.E+07 

Cycles to failure, Nf 

Composite-to-steel lap shear specimens were 

used to evaluate epoxy and polyurethane adhesives 

and included Pliogrip 7779 polyurethane, Betamate 4601 and Betamate 2096 epoxy, Hysol E-20HP 

epoxy, and 3M 2214 aluminum-filled epoxy. Cohesive failures were obtained with specimens preheated slightly (35°C–65°C) to enhance adhesive 

flow and wet-out of the steel and composite surfaces. “Room temperature” curing adhesives also 

responded well to the application of some heat to 

ensure full cure of the adhesive. Lap shear strengths 

for these type specimens are on the order of 20– 

24 MPa. 

% Ultimate Break Load 

Bolted joint -Air Bolted joint - Distilled water Bolted joint - 5 wt.% Salt water Bolted/adhesive joint - 5 wt.% Salt water 

Figure 6. Environmental fatigue test results for bolted 3­

D glass fiber–reinforced epoxy composite/steel lap shear 

specimens. Modeling 0.20 Adhesively bonded joints can often be designed 0.18 to be stronger, stiffer, and lighter than their me­ 0.16 chanical counterparts in many applications. How­ 0.14 

ever, constraints on the design envelope, 0.12 

% Weight Change 0.10 

0.08 

0.06 

Distilled water, sample no.'s 1, 2 and 3 0.04 

5 wt.% NaCl/distilled water, sample no.'s 1, 2 and 3 0.02 Car wash solution, sample no.'s 1, 2 and 3 

0.00 

0 100 200 300 400 500 600 700 800 

Time (h) 

Figure 7. Weight gain versus immersion time for 3-D 

manufacturing processes and the cost of a final 

product often prevent designers from considering 

adhesively bonded joints. 

Bonded joints are susceptible to peel stresses, 

and care must be taken to eliminate and/or minimize 

peel loading near the edges of the joint. Design envelope constraints may preclude the design modifications necessary to minimize peel stresses, thereby 

eliminating bonded joints from consideration. Hartglass fiber–reinforced epoxy composite coupons. 

through the machine edges and hole perimeter when 

Smith analyzed single-lap joints and concluded that 

joints with thin adherends and large overlap areas 

exhibited failure of the adherends with a high joint the coupons are in a static (un-stressed) condition. efficiency. Joints with thicker adherends failed due 

Quasi-static tensile tests were conducted with to peel or shear stresses, and the joint efficiency was 

sets of bolted composite/steel lap shear specimens significantly lower. Joint efficiency with thick adprepared from 3-D glass fiber–reinforced composite herends is possible using scarf or stepped-lap joint coupons that had been immersed in distilled water techniques. However, these methods significantly and 5 wt % salt water for 8 days at room tempera- increase the manufacturing cost and complexity of 

ture. No reduction in joint tensile break loads was the joint. For these reasons, adhesive bonding of detected with either group of specimens. 

Structural Adhesives 

Data presented in the FY 2005 Annual Progress 

thick adherends is generally not considered a viable 

option in automotive and heavy truck applications. 

Mechanical (bolted) joints have several advantages that make them attractive for heavy vehicle 

applications. Unlike bonded joints, bolted joints can Report showed that the fatigue life of bolted com­ withstand significant peel stresses, and loads can be posite/steel assemblies can be enhanced by the additransferred more efficiently through thick compo­ tion of a structural adhesive to the joint. In this nents using simple joint geometries such as single- effort, commercially available structural adhesives 

were and double-lap configurations. In addition, bolted evaluated for their ability to bond well with 

joints can be easily inspected, repaired, require both 3-D glass fiber–reinforced composite and steel 

minimal surface preparation prior to assembly, and substrates, and for ease of processing. have lower production costs. 

75 
FY 2006 Progress Report 

Two load-transfer modes can be idealized for a 

bolted joint loaded in shear such as the double-lap 

specimen. These are hole loading and contact surface loading. The dominant mode type is primarily 

dependent on the bolt preload and is affected by material properties and joint geometry. The holeloading mode occurs when the preload is low and 

components are allowed to slip under load. This 

scenario is not desirable in heavy truck applications 

because vibration and reversal of loads would result 

in movement of the components, quickly resulting in 

material wear near the bolt due to friction. Higher 

bolt preload results in load transfer through friction 

at the component (contact) surfaces. “Failure” of 

these joints occurs when either the interfaces slip or 

the material incurs damage. The ultimate load is 

strongly dependent on the force and contact area. 

Finite element analyses (FEA) were conducted 

for the case of a symmetric double-lap joint consisting of a 3-mm-thick 3-D glass fiber–reinforced 

vinyl-ester composite joined to two steel plates of 3, 

5, 7 and 11-mm thickness via a 12-mm (1⁄2-in.) bolt. 

The joint assembly prepared with the 3-mm thick 

steel plates is a representation of an actual setup 

tested in this project and described earlier in this 

report. The detailed geometry of the joint was idealized and only one-half of the setup was analyzed due 

to its symmetry (Figure 8). 

Figure 8. Idealized representation of double-lap 

joint. 

Preloads (i.e., the clamping force applied to the 

assembly after tightening the bolt) of 0.44, 2.22, 

4.45, 8.90, 13.34, and 22.24 kN (100, 500, 1000, 

2000, 3000, and 5000 lb-ft) were applied to each 

geometry setup for a total of 24 analyses performed. 

The objective was to inspect the contact area between the components as a function of the thickness 

of the steel plates and the applied preload. VonHigh Strength Weight Reduction Materials 

Mises stresses in two cross sections are depicted in 

Figure 9. 

Figure 9. Graphical representation of Von– 

Mises stresses due to preloads. 

Figure 10 shows the limit cases of the joint. The 

conclusion is that while the magnitude of the contact 

pressure changes as a function of applied preload, 

the contact area remains virtually unaffected for 

thinner steel plates and increases only modestly for 

thicker substrates. The contact area can be assumed 

constant and equal to the area of the washer inserted 

between the plates and the bolt or nut for the case of 

the 3-mm-thick steel plates (Figure 10.-a). For 

thicker steel plates, the contact area is greater and 

increases with applied preload as illustrated in Figure 10b. 

Summary 

This project is a 4 year research effort focused 

on the development of technically robust and economically attractive joining techniques to overcome 

the technical issues associated with joining composite chassis components in heavy vehicles. This report discusses the FY 2006 research activities. 

A variable amplitude fatigue testing has been 

conducted with 3-D glass fiber–reinforced epoxy 

composites. An exploratory study was initiated with 

a commercially available carbon fiber–reinforced, 

cross-ply laminate to investigate the fundamental 

relationships between energy dissipation and damage accumulation, and their utility in making fatigue 

life predictions. 

Loss of bolt preload due to through-thickness 

creep of the composite has been monitored for composite/steel joints under static and fatigue loading, 

76 
High Strength Weight Reduction Materials 

(a) 3-mm-thick steel plates. 

(b) 11-mm-thick steel plates 

Figure 10. Contact area and pressure for (a) 3-mm-thick and (b) 11­

mm-thick steel plates with preloads of, respectively, 0.44 and 

22.24 kN (100 and 5000 lb-ft) 

FY 2006 Progress Report 

and commercially available structural adhesives 

were evaluated for their ability to bond well with 

3-D glass fiber–reinforced composite and steel substrates as well as for ease of processing. Environmental fatigue testing has been conducted to 

evaluate the durability of bolted 3-D glass fiber reinforced composite/steel joints with exposure to air, 

water, and salt water environments. 

Finite element analyses were conducted for the 

case of a symmetric double lap joint consisting of a 

3-D glass fiber–reinforced composite joined to two 

steel plates to inspect the contact area between the 

components as a function of the thickness of the 

steel plates and the applied preload. 

Acknowledgements 

The principal investigators would like to acknowledge the valuable contributions of the members of the research team, including Don Erdman, 

Vlastimil Kunc, and Tom Battiste of ORNL. 

77
FY 2006 Progress Report High Strength Weight Reduction Materials 

78 
High Strength Weight Reduction Materials FY 2006 Progress Report 

C. Friction-Stir-Joined Aluminum Sheet Materials for Heavy Vehicle Cab 

Structures 

Principal Investigator: Glenn J. Grant 

Pacific Northwest National Laboratory 

P.O. Box 999, Richland, WA 99352-0999 

(509) 375-6890; fax: (509) 375-4448; e-mail: glenn.grant@pnl.gov 

Chief Scientist: James J. Eberhardt 

(202) 586-9837; fax: (202) 587-2476; e-mail: james.eberhardt@ee.doe.gov 

Field Technical Manager: Philip S. Sklad 

(865) 574-5069; fax:(865) 574-6098; e-mail:skladps@ornl.gov 

Participants: 

Elizabeth Stephens, Karl Mattlin, and Mike Dahl, Pacific Northwest National Laboratory 

Scott Wazny and Brian Fulbright, Freightliner LLC 

Douglas Waldron and Keith McTernan, Advanced Joining Technologies, Inc. 

Tom Sanor, Magna Drive Automotive 

Alcoa 

Contractor: Pacific Northwest National Laboratory 

Contract No.: DE-AC06-76RL01830 

Objective 

• Develop and deploy friction stir joining (FSJ) as a weight and cost-saving manufacturing technology for heavy 

vehicle cab structures 

Approach 

• Demonstrate the use of aluminum tailor-welded blanks (TWBs) for heavy vehicle applications. 

• Develop weld process parameters to lower manufacturing cost, increase joint reliability, and explore dissimilar 

material blanks that can afford additional weight savings. 

• Develop and prototype lightweight cab-in-white structures, including door inner liners, door opening panels, 

back- wall, and floor structures for Class 7–8 trucks using FSJ technologies. 

• Address manufacturing issues to lower blank fabrication cost and other barriers to implementation of aluminum 

TWBs. 

Recent Accomplishments 

• Began conducting a design of experiments on tool development and process design for high-speed welding 

(+150 in. per min) to lower the cost of the initial blanking operation. 

• Developed three tool designs for high-speed welding of dissimilar aluminum alloys of different thicknesses 

(5182-O and 6111-T4 aluminum alloys). 

• Developed two tool designs for high-speed welding of dissimilar aluminum alloys of similar thicknesses. 

79 
FY 2006 Progress Report High Strength Weight Reduction Materials 

• Completed setup of digital image correlation for use with limited dome height tests to track strain localization 

and strain variation of welded parts during forming. 

• Characterized and completed strength and formability tests for the 5182/6111 TWBs fabricated for high-speed 

welding. 

Future Direction 

• Continue with the weld process and tool development for dissimilar alloy combinations and hard alloys (6xxx, 

5xxx H3x, etc.) that would result in greater cost and weight savings. 

• Develop and optimize the welding process parameters of dissimilar aluminum alloys of different thicknesses 

(5182-O and 6111 aluminum alloys). 

• Focus on FSJ as a low-cost enabler of lightweight, low-cost vehicle assembly (non-TWB). 

Introduction 

This work is a collaborative effort between Pacific Northwest National Laboratory (PNNL), 

Freightliner, LLC, Advanced Joining Technologies, 

Inc. (AJT), Magna Drive Automotive, and Alcoa. 

This project aims to develop and deploy FSJ as a 

weight- and cost-saving manufacturing technology 

for heavy vehicle cab structures. To date, the project 

has focused on (1) developing the appropriate FSJ 

process parameters to create high-reliability joints 

that survive stamping operations; (2) prototyping 

full cabs using FSJ technologies and testing for durability and crash performance; and (3) addressing 

manufacturing issues, including lowering blank fabrication cost, and other barriers to implementing the 

technology. The challenge is to develop stamped 

panels that can meet the unique strength and durability requirements of heavy vehicle cab structures. 

Aluminum TWBs consist of multiple-thickness 

and multiple-alloy sheet materials welded together 

into a single, variable-thickness blank. Figure 1 

shows a typical fusion-welded TWB before and after 

a stamping application. A TWB is assembled as a 

series of flat sheets joined together, which are then 

submitted to a stamping process. The technology 

allows production of a weight-optimized, variablethickness vehicle body component. TWB technology 

gives automotive and truck designers the ability to 

selectively vary body panel thickness to optimize the 

use of material. Successful use of the technology 

ultimately results in reducing vehicle weight without 

compromising final strength, stiffness, and durability. The manufacture of TWBs and their application 

in body panels requires that the weld material deform under biaxial loading during sheet metal 

stamping. The deformation of weld materials and 

(a) 

(b) 

Photos courtesy of Reynolds Metals Company and Ogihara America Corp. 

Figure 1. (a) A TWB viewed in the as-welded condition, ready for submission to the stamping process. 

The shape shown is typical for a door inner stamping 

operation. (b) Aluminum TWB after stamping to produce a door inner panel. 

their limits of formability are important aspects of 

TWB technology. 

The primary challenge of using aluminum 

TWBs in the past has been the relatively low quality 

of aluminum fusion welds. It often results in premature fracture or a lack of reliability of weld materials 

80 
High Strength Weight Reduction Materials 

during stamping. Improving aluminum weld quality, 

understanding and describing the formability of the 

weld region, and predicting its formability are the 

primary technical challenges in using aluminum 

TWBs. 

FSJ is a revolutionary joining technology that 

employs severe plastic deformation to create solid 

state joints between wide varieties of different materials. Invented by TWI, Ltd., about 12 years ago, 

FSJ is capable of producing aluminum alloy welds 

as good as (or significantly better than) fusion welds 

in terms of joint efficiency, mechanical properties, 

and environmental robustness. The advantage of 

using this solid state joining technique is the ability 

to avoid liquid metal during joining. Aluminum has 

low molten viscosity, high reflectivity, and a relatively high propensity to form internal porosity during joining because of the high solubility of 

hydrogen in liquid aluminum. 

FSJ also can eliminate hot cracking and minimize heat-affected-zone (HAZ) issues in 6000-series 

heat-treatable aluminum alloys. Avoiding the liquid 

phase of the materials also avoids the formation of 

various large eutectic constituent particles or other 

undesirable intermetallic particles that develop in 

particular alloys or with certain types of weld or materials contamination. 

The use of FSJ also better facilitates welding of 

heat-treatable materials, since the HAZ is normally 

smaller or less pronounced than with fusion welding. 

The use of solid state joining enables a variety of 

types of dissimilar aluminum alloy joining not normally possible using fusion welding methods. The 

goal of producing a highly formable weld joint may 

be achievable with friction stir welding because of 

the wide range of weld heat, plastic work, and weld 

metal grain-size manipulation that is possible. 

High-Speed Welding 

Although the industrial partners consider the 

program research and engineering a success, there 

are still barriers to implementation of the FSJ technology for aluminum TWBs. While many material 

combinations were successfully stamped, others that 

would have had stronger manufacturing arguments 

(cost and weight savings) were not successfully 

stamped (i.e. 6xxx and 5xxx H3x alloys). In addition, the cost models were close but were still below 

cost parity with monolithic sheet stampings. A lowcost, fast process will be the driver for implementaFY 2006 Progress Report 

tion of the technology. Joint reliability and speed of 

blank fabrication are two barriers to overcome. 

To overcome the blank cost, high-speed welding, >150 in. per min (ipm), of 5182-O and 6111-T4 

TWBs is being investigated (2 and 1.5mm thick, 

respectively). Currently, a design of experiments 

(DOE) on tool development and process design is 

being performed. Three tool designs were developed 

for high-speed welding of the dissimilar aluminum 

alloys of different thicknesses. A 2-level factorial 

design was chosen to study the joint effect of process parameters on formability. Two rounds of experimental testing will occur. The first round of 

testing is designed to hone in on one tool design for 

high-speed welding; the second round of testing is to 

continue with optimizing the welding process parameters with the selected tool. 

For the first round of testing, the two factors 

considered for the 2-level factorial were travel speed 

and tool rotation speed. A low-level and high-level 

speed for each factor were determined. For tool 

travel speed, a low level of 150 ipm and a high level 

of 300 ipm were chosen. For tool rotation speed, a 

low level of 2000 rotations per minute (rpm) and a 

high level of 3500 rpm were chosen. 

These levels were chosen for the 2-level factorial because previous work indicated these values as 

the feasible ranges for these materials. Table 1 is 

representative of the DOE matrix created. A 5182-

O/6111-T4 TWB panel was friction stir welded for 

each run (at the determined process parameters) in 

the DOE matrix. The matrix consists of six weld 

runs, and replicates of each run will be performed. 

Table 1. DOE matrix developed for tool design 

Travel Rotation 

speed speed 

Run Matrix (ipm) (rpm) 

1 –1 +1 → 150 3500

2 –1 –1 → 150 2000

3 +1 –1 → 300 2000

4 +1 +1 → 300 3500

5 0 0 → 225 2750

6 0 0 → 225 2750

In the DOE, each weld run (including replicates) 

is treated as an independent experiment. Each weld 

run must be conducted randomly to rule out any 

variability and influences. All material alloys used 

for the weld runs come from the same source (i.e., 

81 
FY 2006 Progress Report 

same supplier, same sheet of metal). We will be able 

to ensure uniformity among the materials and be 

able to determine that any differences observed in 

formability performance from subsequent testing are 

due to the tool and not to the material used. 

Test coupons were prepared from the TWB panels made from each weld run and were subsequently 

tested for strength and formability. Tensile tests and 

limited dome height (LDH) tests were performed. 

Further investigation of the optimum joining process 

parameters will occur in the second round of testing, 

in which a second DOE matrix will be created to 

establish the optimum weld conditions. Test coupons will be prepared from the welded panels and 

subsequently tested and characterized to determine 

the set of parameters that produces a TWB with the 

best formability. 

Forming Strain Characterization Using 

Digital Image Correlation 

During this reporting period, digital image correlation (DIC) has been applied for use with limited 

dome height formability tests to overcome the barrier of joint reliability. DIC is a data analysis method 

that uses an algorithm to analyze digital image data 

taken when a sample is subjected to mechanical 

strain. Displacement and strain are measured. This 

technique uses white-light speckle correlation, 

where two similarly speckled images captured by a 

video camera represent the state of the object before 

and after deformation. Consecutive images are captured during testing, and the image correlation will 

register a change in surface characteristics as the 

specimen is affected by stresses imposed upon it. 

The actual object movement is measured and the 

Lagragian strain tensor is available at every point on 

the surface. 

This technique will help us further understand 

strain localization in a non-homogenous gage such 

as a weld. We will be able to track strain localization 

and strain variation of the welded parts during forming. These incremental strains acquired during forming will be entered into finite element analysis 

(FEA) models. This technique will allow us to comprehend the stresses required for forming. The use of 

FEA models will allow us to re-create the loading or 

stress paths that occur during deformation. Figure 2 

illustrates the major strain observed in a monolithic 

6111 sheet during LDH testing. 

High Strength Weight Reduction Materials 

Figure 2. An illustration of the major strain characterization of a 6111-T4 monolithic sheet during LDH testing. 

Note that image was taken prior to reaching peak height. 

Tool Designs and Blank Fabrication 

Maximum weld traverse speed is very much a 

function of tool design. Three tool designs were developed for high-speed welding of the dissimilar 

aluminum alloys of different thicknesses to address 

the speed of blank fabrication. Following is a description of each tool design used for the DOE. 

Tool A: AJT proprietary tool 

Tool B: Threaded, three-flat pin tool with flats 

equally spaced, convex scrolled shoulder 

Tool C: Threaded, three-flat pin tool with flats 

equally spaced, flat scrolled shoulder 

All blanks were fabricated with a lead angle of 

one degree in the direction of weld travel. 

Formability and Strength Performance of 

High-Speed Welds 

Limited dome height testing has shown that 

welded blanks with the highest dome heights can be 

correlated with those that have minimal property 

gradients across the weld zone. Minimal property 

gradients delay the onset of necking instability. 

Changes in process parameters will change the local 

mechanical properties and provide a way to “customize” the weld to the parent sheet. Picking process 

parameters that match flow stresses between the 

sheet and the weld zone may be a strategy to increase stamping performance. 

82 
High Strength Weight Reduction Materials 

In the limited dome height evaluations, two 

specimens per weld run and tool design in the DOE 

were tested. Similar trends were observed. All the 

specimens tested, regardless of tool design and weld 

travel speed, failed at low loads and “unzipped” on 

the weld center-line. The weld splitting indicates a 

poorly consolidated weld. Based on the DIC analysis, the strain accumulated in the parent sheets next 

to the weld prior to failure ranged from 3–6% depending on the weld process parameters and tool 

design. Figure 3 is a representative photo of a 

specimen during testing illustrating the failure mode 

observed. Figure 4 is a comparison of the weld 

speed versus failure load of process parameter data 

from earlier studies and the recent tool design limited dome height test data. 

Figure 3. Representative photo of a limited dome height 

specimen during testing. Unzipping of the weld at the 

weld centerline can be observed. 

Weld Speed vs Failure Load 

TWB 5182-O 2mm to 6111-T4 1.5mm 

Failure Load (lb) 10,000

9,000

8,000

7,000

6,000

5,000

4,000

3,000

2,000

1,000

0 

0 50 100 150 200 250 300 350 

Previous Data 

Tool A 

Tool B 

Tool C 

Weld Speed (ipm) 

Figure 4. Comparison of the weld speed versus failure 

load of process parameter data from earlier studies and 

the recent tool design limited dome height test data. 

FY 2006 Progress Report 

For the tensile tests, ASTM E-8 sub-size specimens were used. Four specimens were tested for 

each weld run and tool design in the DOE. The peak 

load strengths varied among the specimens depending on the tool and process parameters. All the 

specimens fabricated from Tool C, regardless of 

welding speed, failed in the weld centerline. Most 

specimens fabricated from Tool A, except three, also 

failed in the weld centerline. Two of the four specimens fabricated at 150 ipm and 2000 rpm and one of 

the four specimens fabricated at 225 ipm and 2750 

rpm failed in the 6111 thin sheet material near the 

HAZ. As with Tool A, most specimens fabricated 

from Tool B, except three, failed in the weld centerline. Two of the four specimens fabricated at 

150 ipm and 3500 rpm and one of the four specimens fabricated at 225 ipm and 2750 rpm failed in 

the 6111 thin sheet material near the HAZ. Figure 5 

is a representative photo of the two failure modes 

observed among the tensile specimens. 

(a) (b) 

Figure 5. Photo representative of the specimen failure 

modes observed after tensile testing: (a) in the 6111-T4 

thin sheet material near the HAZ and (b) unzipping in the 

weld centerline. 

Overall, the formability and strength results of 

the high-speed welds in this current work were poor. 

The results are significantly below those from previous work, as illustrated in Figure 4. It is speculated 

that a weld defect was generated during welding of 

the blanks (this will later be verified through microstructural analysis) and contributed to the “unzipping” of the welds observed during testing. In 

addition, the welding setup conditions were not optimum for Tool B and Tool C. These two tools were 

designed with longer pin lengths to ensure full con83 
FY 2006 Progress Report 

solidation of the weld and were designed to be used 

with a transverse tilt (a rake angle perpendicular to 

weld travel) to accommodate the dissimilar sheet 

thicknesses. The pin lengths were also shortened 

during fabrication of the blanks. 

To continue with the investigation of high-speed 

welding, new welded specimens will be fabricated 

with the same three tool designs; however they will 

implementing a transverse tilt with no modification 

to the pin lengths. The DOE test matrix will be repeated. All results from the limited dome height and 

tensile tests will be fed into the DOE to hone in on 

one tool design for high-speed welding. Further investigation of the optimum joining process parameters will occur in a second round of testing, in which 

a second DOE matrix will be created to establish the 

optimum weld conditions. 

High Strength Weight Reduction Materials 

Conclusions 

From this investigation, the following conclusions were derived: 

• FSJ is suitable for fabricating aluminum TWBs. 

• The FSJ process is “customizable” for mechanical properties, and the weld process parameters 

can be optimized for formability. 

• Further investigation of tool design and weld 

process development is needed to determine optimum joining process parameters for highspeed welding of 5182-O /6111-T4 TWBs. 

Publications 

G. J. Grant, R. W. Davies, E. V. Stephens, et al., 

“The Formability of Friction Stir Welds in Automotive Stamping Environments,” SAE Transactions 

Journal of Materials & Manufacturing 114(5), 619– 

629 (2006). 

84 
High Strength Weight Reduction Materials FY 2006 Annual Progress Report 

D. Friction Stir Joining and Processing of Advanced Materials including MMCs 

Principal Investigator: Glenn J. Grant 

Pacific Northwest National Laboratory 

P.O. Box 999, 902 Battelle Blvd., Richland, WA 99352 

(509) 375-6890; fax: (509)375-44; e-mail: glenn.grant@pnl.gov 

Project Manager: Moe Khaleel 

Pacific Northwest National Laboratory 

P.O. Box 999, 902 Battelle Blvd., Richland, WA 99352 

(509) 375-2438; fax: (509) 375-6631; e-mail: moe.khaleel@pnl.gov 

Chief Scientist: James J. Eberhardt 

(202) 586-9837; fax: (202) 587-2476; e-mail:james.eberhardt@ee.doe.gov 

Field Technical Manager: Philip S. Sklad 

(865) 574-5069; fax: (865) 574-6098; e-mail: skladps@ornl.gov 

Participants: 

Bill Arbegast, Stan Howard, Casey Allen 

Advanced Materials Processing Center 

South Dakota School of Mines and Technology 

Rapid City, S.D. 

Contractor: Pacific Northwest National Laboratory 

Contract No.: DE-AC06-76RL01830 

Objective 

• Develop friction stir joining and processing (FSJ/P) to enable widespread use of lightweight, advanced materials in heavy vehicle manufacturing. 

Approach 

• Develop new technologies within the broad topic area of FSJ/P that facilitate lightweight manufacturing methods and new materials for use in heavy vehicle application. 

• During 2006 the project focused on two primary areas: 

• Using FSJ to join sheet materials for superplastic forming applications 

• Developing fundamental process models using a new approach: smooth particle hydrodynamics (SPH) 

Accomplishments 

• Developed FSJ weld metal with superplastic properties to facilitate large, multisheet SPF or integrally stiffened 

panel applications for heavy vehicle cab structures. 

• Fabricated proof of concept multisheet panels joined together by FSJ and SPF formed into stiffened structural 

panels. 

• Developed unique computational code based on SPH to predict material flow, mixing and defect formation in 

FSJ/P. 

• Collaborated with researchers at the South Dakota School of Mines and Technology (SDSMT)in a numerical 

modeling study of the strength and dampening characteristics of fabricated multisheet panels. 

85 
FY 2006 Annual Progress Report High Strength Weight Reduction Materials 

Future Direction 

• Develop weld process parameters that enhance superplasticity in joined assemblies to allow for part consolidation in lightweight SPF cab structures. 

• Test multisheet and integrally stiffened panel for strength, modulus, and other mechanical performance metrics. 

• Parallelize SPH computational code to run on supercomputers to improve the ability of the model to define fine 

features of flow, defect formation, and tool designs in friction stir joined or processed material. 

Introduction 

Friction Stir Joining 

One of the key strategies for making a vehicle 

energy-efficient is to manufacture it from lighter 

materials. Structural and functional requirements, 

however, lead to a situation in which no single 

lightweight material is appropriate for all applications. A modern, weight-optimized vehicle structure 

is a hybrid of many materials. A critical problem that 

has emerged in the development of these hybrid 

structures is that for many material combinations, 

traditional joining technologies (e.g., fusion welding 

or mechanical fastening) are not appropriate. For 

some highly specialized materials, like aluminum 

metal matrix composites (MMCs), titanium, and advanced high-strength steels, a better joining technology can have significant impact on whether these 

materials have a role in future vehicle structures. 

In the past 15 years, the new joining technology, 

FSJ, has emerged and has the potential to join many 

lightweight materials. Invented by TWI, Ltd., FSJ is 

a solid-state process that employs severe plastic deformation to create joints between wide varieties of 

different materials. A typical FSJ butt joint is depicted in Figure 1. The weld is created by clamping 

the materials to be joined and plunging a spinning 

tool into the surface. The spinning tool is then translated down the joint line, leaving behind a weld zone 

characterized by a fine-grained, dynamically recrystallized microstructure. Typically, the tool is spun at 

400 to 2000 rotations per minute (rpm) and translated down the joint line at a rate of 4 to 300 in./min 

(ipm), depending on tool design, base material, and 

thickness. As the tool rotates and translates, complex 

flow patterns develop in the base material that create 

an intimate mixing of materials from both sides of 

the weld. Heat input during plastic deformation generally creates a temperature in the weld between 0.6 

and 0.8 of the absolute melting temperature so that 

no liquid phase is generated. 

Figure 1. Friction stir joining and processing are accomplished by plunging a spinning tool into a material 

and translating the tool across the surface to form either 

a joint or a surface processed region (Source: TWI, 

Ltd.). 

FSJ is capable of producing aluminum and magnesium alloy welds as good as, or better than, fusion 

welds in terms of joint efficiency, mechanical properties, and environmental robustness. A significant 

advantage of the process, for application to hybrid 

structures, is that since there is no melting during the 

process, a large variety of dissimilar material joints 

are possible, including dissimilar aluminum and 

magnesium joints that are not possible with conventional fusion welding. 

In the past 10 years, FSJ has been shown to be a 

commercially important, energy-efficient, and environmentally friendly process for joining aluminum. 

However, there are many opportunities for other 

higher-strength lightweight materials to be considered if good joining technologies existed for these 

materials as well. The objective of this project is to 

investigate how FSJ can be applied to advanced materials including AL-MMCs, titanium, steels, cast 

iron, superplastic materials, and materials that display graded structures with unique surface properties. 

86 
High Strength Weight Reduction Materials FY 2006 Annual Progress Report 

Approach 

The basic objective of this project is to investigate and develop FSJ and FSP as viable industrial 

techniques for advanced lightweight materials. During 2006, the project concentrated on two main task 

areas: (1) development of FSJ/P as a process to join 

sheets that will be subjected to superplastic forming, 

and (2) development of numerical, physics-based 

modeling strategies to help define the fundamental 

process and predict process conditions. 

The primary task in 2006 was the application of 

FSJ to superplastic forming environments. Numerous opportunities exist to use this manufacturing 

technology to reduce weight on heavy vehicles. One 

barrier to SPF manufacturing is that hang-on truck 

components are complex and often very large— 

larger than the standard coil widths. If sheets are 

joined together to make a large part by fusion welding, the weld region does not deform superplastically 

because of the coarse microstructure of fusion weld 

metal. FSJ produces a weld region that is formable 

by SPF and may be an enabler for complex or large 

structures that are joined, then later hot gas formed 

in single operations to form cab structures and panels. 

The second task investigated was continued development of a numerical modeling process using 

SPH. This modeling approach has not been previously applied to FSJ/P, but our preliminary work 

suggests that it may be able to provide significant 

insight into the fundamental nature of the heat generation and material flow that occurs during the 

FSJ/P process. 

Results 

Task 1 — FSJ/P for SPF manufacturing 

Superplastic forming (SPF) is a lightweight 

manufacturing technology that has shown to be cost 

effective and produce weight-optimized structures 

through reduction in part count and through the use 

of lightweight materials. 

One important factor that is relevant to the particular problems of heavy vehicle SPF is that truck 

parts are large. The parts are modulus (stiffness) 

driven, not strength driven, because of aero flutter 

and shear size. Big parts need stiffeners. This requires a complex assembly. One cost effective and 

lightweight option is to join multiple flat sheets together, then use SPF to join the entire multisheet 

pack together in one operation. This incorporates the 

stiffener into the assembly. The problem with this 

strategy revolves around the method chosen to join 

the sheets together. If the sheets are fusion welded, 

the microstructure of the fusion weld region will not 

deform superplastically. Figure 2 shows the coarse 

microstructure of a fusion welded aluminum sheet. 

When these joints are subjected to superplastic 

forming, the joint remains undeformed. These undeformed areas also have very poor postforming mechanical properties. In contrast, FSJ leaves behind a 

weld microstructure that is highly deformable at SPF 

conditions. 

(a) 

(b) 

(c) 

Figure 2. (a) Weld nugget region in a butt weld between 

aluminum sheet materials showing coarse, cast microstructure of a fusion weld. (b) Transverse tensile specimen. (c) 

Transverse tensile test at SPF conditions shows the undeformed nature of the fusion weld metal. 

This project is investigating the particular weld 

process parameters and part geometries that could 

produce an integrally stiffened panel for a heavy 

vehicle cab application. The work to date involves 

fabricating multisheet assemblies as shown in Figure 3. These assemblies have been SPF formed into 

structural panels in three configurations, “egg crate,” 

“hat stiffened,” and “doughnut.” The purpose of the 

test forms is to discover the correct FSJ process parameters and the best SPF deformation possible in 

these geometries. The panels are made using both 

linear FSJ and friction stir spot welding (FSSW) 

87 
FY 2006 Annual Progress Report High Strength Weight Reduction Materials 

Figure 3. Geometries of test panels fabricated for this 

program 

(Figures 4 and 5). The FSSW joints are of particular 

interest because they are made using a new process 

that does not leave an exit hole after the spot joint is 

made. This process, called the “refill” method was 

invented by GKSS (Riftec) and the panels were fabricated at the facilities of our project partner, 

SDSMT. 

Figure 4. Friction stir spot welded three-sheet multisheet 

pack prior to SPF testing. This configuration is designed 

to produce the corrugated structure. Alternating spot 

welds are made from each side of the multisheet pack 

using the “Refill” (GKSS-Riftec) FSSW method at the 

SDSMT. 

Structural panels like those shown in Figures 6, 7, 

and 8 were successfully formed by SPF at 530oC, 

both with and without die constraint. Offset linear 

friction stir joints on the multisheet pack produce a 

truss structure, and offset friction stir spot welds 

produce an egg crate structure (truss on the diagonal). 

Figure 5. Linear friction stir welded multisheet pack prior 

to SPF testing. This configuration is designed to produce 

the “three sheet with hat stiffener” structure. Alternating 

linear welds are made from each side of the multisheet 

pack using a pin that penetrates only through the upper 

two sheets. 

Figure 6. Structure formed from multisheet packs joined 

together as sheets then subsequently hot-gas formed into a 

lightweight structural panel 

Figure 7. Sheet pack cross section after SPF. This structure 

was fabricated with no constraining die platens. 

A particular advantage of shapes like those 

shown in Figure 8 is that they are hollow with interconnected channels between the outer sheets. This 

shape gives the panel a 3-D structural stiffness, but 

88 
High Strength Weight Reduction Materials FY 2006 Annual Progress Report 

Figure 8. Two side cross section of a friction stir spot 

welded multisheet structure using die platens with a 1- 

in. standoff. 

it also gives it particular properties for dampening, 

vibration, energy absorption, and thermal conductivity. Essentially, the airspace and shape of the structure acts to isolate one side of the sheet from 

another. This allows these panels to have very good 

properties as noise, vibration, harshness (NVH) resistant structures. They also can be used to thermally 

isolate one area from another as is might be required 

for a firewall structure, or used as a crash energy 

absorption structure. The inclusion of space filling 

foams either metallic or polymeric can further enhance these properties. 

To investigate the effects of different internal 

shapes on the static, vibrational, and energy absorbing characteristics, researchers at the facilities of our 

project partners at SDSMT have completed a numerical model of the static stresses and modal frequencies of the different panels fabricated at PNNL 

(Figure 9). The results of this work illustrate the 

positive advantages that “sandwich” panels can have 

as compared with monolithic sheet. 

Figure 9. Von Mises contour plots of the sandwich 

panels from static analysis. (Source: Karim H. MuciKuchler, et al., 2007) 

Task 2 — FSJ/P Numerical Modeling 

Work continued in FY 2006 on a new modeling 

approach to investigate the fundamental flow and 

heat generation properties of the friction stir process. 

The industry has struggled with modeling this process for 10 years because the high strains and high 

strain rates are very difficult to capture in conventional modeling approaches. Conventional modeling 

techniques such as finite element modeling and 

computational fluid dynamics suffer from many issues, mostly surrounding the need for the model to 

track huge variations in strain, temperaturedependent flow stress, strain rate, and heat generation. 

The modeling approach we have started to investigate in this program is a departure from other 

approaches and involves the use of SPH. This is a 

fully Lagrangian particle method, so no grids are 

required. The code solves 2-D and 3-D equations for 

continuity, momentum, and energy balance. The 

model uses a temperature-dependent shear stress 

equation (Johnson Cook formulation) to track 

stress/strain response so material databases are easier to incorporate as input parameters. 

The model currently can explicitly model mixing of materials, predict temperature generation at 

the surface of the tool as well as that due to viscous 

dissipation, and produce temperature and strain histories for any material point in the field. This may 

allow it to make microstructural prediction at different parts of the weld since temperature and strain 

history can be tracked, although this has not as yet 

been implemented in the code. Figure 10 shows 

some example output from the model. 

Fully implemented, the code should be able to 

make predictions on the effects of tool shape and 

changing process parameters and help to provide a 

fundamental physical understanding of the FSJ process. 

Conclusions 

FSJ/P technologies will enable the application of 

many lightweight materials in the next generation of 

transportation systems. It is envisioned that structural multisheet panels can play a key role in making 

vehicles lighter weight. In addition, these panels 

hold opportunities for multiuse structures, such as 

metal or polymeric foam filled sheet packs with 

buckle resistance or internal dampening for better 

NVH performance. 

89 
FY 2006 Annual Progress Report High Strength Weight Reduction Materials 

The results of this work will allow designers to 

anticipate and implement structures that are a hybrid 

of many different materials, facilitate the application 

of lightweight superplastically formed structures, 

and suggest new materials and engineered surfaces 

that can help deliver lighter and more fuel-efficient 

vehicles. 

Figure 10. Model output. Upper two figures show mixing of dissimilar materials using a tool with three flats 

on the pin. (Only the tool pin is shown in this 2-D simulation.) The lower two figures show arrangement of 3-D 

model and a temperature distribution. Highest temperatures are in a region near the pin root away from the pin 

surface, due to heat generated by plastic work and conduction into the pin. 

Presentations/Publications/Patents 

1. Tartakovsky, A., Grant, G., Sun, X., and 

Khaleel, M., “Modeling of Friction Stir Welding 

(FSW) Process with Smooth Particle Hydrodynamics (SPH),” Welding and Joining and Fastening and Friction Stir Welding, SAE Special 

Publication SP-2034, SAE, 2006. 

2. Grant, G.J., Herling, D., Arbegast, W., Allen,C., 

Degan,C., “Superplastic Forming of Aluminum 

Multisheet and Integrally Stiffened Structures 

Fabricated Using Friction Stir Welding and Friction Stir Spot Welding,” (abs) 9th International 

Conference on Superplasticity in Advanced Materials (ICSAM 2006), Chengedu, China, June 

2006. 

3. Grant, G.J., Herling D., “Friction Stir Welding 

to Facilitate Superplastic or Quick Plastic Forming of Complex Automotive Structures,” (abs) 

Society of Automotive Engineers 2006 World 

Congress, Detroit, April 2006. 

4. Tartakovsky A.M., G.J. Grant, X. Sun, S. Hong, 

and M.A. Khaleel, “Smooth Particle Hydrodynamics (SPH) Model For Friction Stir Welding 

(FSW) of Dissimilar Materials.” In 6th International Symposium on Friction Stir Welding, 

Saint-Sauveur, Canada, October 10, 2006. 

PNNL-SA-52321, 2006. 

5. Muci-Kuchler, K.H. S.K. Annamaneni, D.R. 

Herling, and W.J. Arbegast, “Numerical Simulation of the Static and Dynamic Response of Corrugated Sandwich Structures made with Friction 

Stir Welding and Superplastic Forming,” In 

Friction Stir Welding and Processing IV, Eds. 

R.S. Mishra, M.W. Mahoney, T.J. Lienert, and 

K.V. Jata, TMS (The Minerals, Metals and Materials Society), to be published in 2007. 

90 
High Strength Weight Reduction Materials FY 2006 Progress Report 

E. Friction Stir Welding and Processing of Advanced Materials 

Principal Investigator: Zhili Feng 

Oak Ridge National Laboratory 

1 Bethel Valley Road, Oak Ridge, TN 37831 

(865) 576-3797; fax: (865) 574-4928; e-mail: fengz@ornl.gov 

Chief scientist: James J. Eberhardt 

(202) 586-9837; fax: (202) 587-2476; e-mail: james.eberhardt@ee.doe.gov 

Field Technical Manager: Philip S. Sklad 

(865) 574-5069; fax: (865) 574-6098; e-mail: skladps@ornl.gov 

Participants: 

Michael L. Santella, Oak Ridge National Laboratory 

Tsung-Yu Pan, Ford Motor Company 

Contractor: Oak Ridge National Laboratory 

Contract No.: DE-AC05-00OR22725 

Objective 

• Develop the technological basis for friction stir welding and processing (FSW/P) of advanced high-strength 

and lightweight materials for the automotive industry. 

• Gain fundamental understanding of the relationships between workpiece and tool material properties during 

FSW/P. 

• Characterize the mechanical properties and microstructures of joints. 

• Correlate the proprieties and microstructures produced by FSW/P to the process conditions. 

Approach 

• Conduct experimental welding and processing tests on advanced materials using the state-of-the-art FSW/P 

process development system. 

• For particular workpiece materials, select the tool material based primarily on high-temperature strength, wear 

resistance, and chemical compatibility. 

• Evaluate mechanical properties and their correlation with microstructures produced by FSW/P. 

• Develop predictive modeling capability to study the properties as a function of process conditions. 

Accomplishments 

• Completed a study of FSP to improve properties of cast magnesium (Mg) alloy AM60B. 

• Demonstrated the feasibility of friction stir spot welding (FSSW) of Mg alloy AZ31. 

• Completed further process development to improve the properties of FSSW of aluminum (Al) alloy. 

91 
FY 2006 Progress Report High Strength Weight Reduction Materials 

Future Direction 

• Investigate the effects of heat treatment on the fatigue life of FSP cast Al alloys. 

• Develop new approaches for FSSW without an exit hole to improve structural properties of the joint. 

• Carry out FSSW of other lightweight materials and dissimilar welds. 

Introduction 

Friction stir welding (FSW) is a novel solid state 

joining method that was described in detail in previous reports. The uniqueness of the process is creating new opportunities for welding both mature 

alloys and newer alloys, as well as enabling their 

consideration in a wider range of applications. 

The use of FSW for many wrought Al plate alloys has been a great success. Nevertheless, the 

technology is still rapidly evolving. There is tremendous interest in and efforts to extend the technology 

to welding a wide range of both high-performance 

lightweight alloys and high-melting-temperature 

alloys. In addition, novel applications of the process 

other than welding are emerging. For example, it can 

be used to thermomechanically process (FSP) a material for microstructure refinement and property 

enhancements. 

This program aims at advancing FSW/P technology to promote the increased use and adoption of 

high-strength lightweight materials in automotive 

structures, particularly heavy vehicles. It supports 

the goals of the Office of Heavy Vehicle Technology to increase fuel efficiency and reduce emissions 

of heavy trucks by weight reduction without sacrificing strength and functionality. 

In FY 2006, we continued studies on FSP of cast 

Mg alloys and FSSW of wrought Mg alloys and Al 

alloys. Part of the project involves a cooperative research and development agreement (CRADA) with 

Ford Motor Company on FSW of wrought and cast 

Al and Mg alloys for vehicle structure applications. 

An effort related specifically to tool material development is also continuing. 

Feasibility of Spot Friction Welding AZ31 

Magnesium 

Magnesium alloys are receiving a great deal of 

attention in the automotive industry because of the 

weight-saving potential they offer compared with 

both steel and Al alloys. Most current Mg parts are 

produced from cast Mg alloys, such as AZ91 and 

AM60; however, wrought Mg alloys such as AZ31, 

AZ61, AZ80 and ZK60 are beginning to draw attention for body and chassis parts.1

. Among these 

wrought alloys, AZ31 is the most common Mg alloy 

available in sheet form. They are produced by direct-chill casting, followed by hot rolling, or by a 

twin-roll continuous casting process. 

There have been relatively few studies on the joining of wrought Mg alloys. Fusion welding—such as 

gas metal arc welding, gas tungsten arc welding, 

resistance welding, laser welding, and e-beam welding—can be used with wrought Mg alloys, but it can 

cause the formation of a brittle inter-metallic phase 

(Mg17Al12) in the welds,1

 causing reduced weld 

strength. 

In comparison, FSW offers a significant quality 

advantage over the conventional fusion welding 

processes for Mg alloys. Fusion welding requires 

parent metal to melt and solidify, causing potential 

defects such as voids, trapped gas, and solidification 

cracking. FSW does not require melting of the parent metal and thus eliminates the formation of these 

defects. 

Experimental 

Magnesium alloy AZ31B-H24 sheets, 1.58 mm 

thick, were obtained from Magnesium Elektron N.A. 

The nominal composition was 3.0 wt.% Al, 1.0 

wt.% Zn, 0.2 wt.% Mn, and balance Mg. The sheets 

were heated to 0 temper at 412°C for 10 minutes 

before spot friction welding experiments. Sheets 

were cut to 100×30 mm coupons. The spot-frictionwelded specimens were prepared in lap-shear configuration with 30-mm overlap. Two different sample surface conditions were prepared. One was asreceived with a dark-gray oxide layer on the surface. 

The second group was cleaned and wire-brushed to 

remove most of the surface oxides. 

The spot friction welding process was carried 

out using an MTS Systems FSW system at Oak 

Ridge National Laboratory. The SFW tool was a 

fixed-pin probe-type tool with a 15-mm diameter at 

the tool shoulder. The 15-mm-diameter spot friction 

92 
High Strength Weight Reduction Materials 

welding tool was chosen because it was close to the 

recommended electrode diameter (15.9 mm) for resistance spot welding of 1.6-mm-thick AZ31B.2,3 

The process used a displacement-control algorithm 

with an insertion depth of either 2.4 or 2.8 mm and 

four different rotational speeds—500, 1,000, 1,500 

and 2,000 rpm. The process time was fixed at 2 s for 

each joint, which was considered to be typical of a 

production operation. Tensile tests of lap shear coupons were conducted on an Instron mechanical testing machine with a crosshead speed of 12.5 

mm/min. Three replicates were produced for each 

joining condition. 

Results and Discussion 

Lap Shear Strengths 

Maximum tensile loads, identified as lap shear 

strengths, of spot-friction-welded AZ31 versus tool 

rotational speed are presented in Figure 1 with the 

standard deviation associated with each averaged 

data point. In general, the specimens with surface 

oxides removed had higher strength than the asreceived specimens. The highest lap shear strength 

achieved was 4.75 kN with oxide-removed 

specimens at 2.4-mm insertion depth and 1000 rpm. 

Specimens with surface oxide cleaned showed 

failure occurring as circumferential cracking around 

the weld button, whereas for the as-received 

specimens, bond failure occurred between the two 

sheets at the spot weld location (interfacial), as 

shown in Figure 2. 

The lap shear strength was also dependent on the 

rotational speed and the insertion depth. Increasing 

the rotational speed for the specimens with surface 

oxide removed from 1000 to 2000 rpm reduced the 

lap shear strength from 4.75 to 3.9 kN. The low 

rotational speed of 500 rpm produced the weakest 

spot-friction-welded joints. The as-received 

specimens produced with a 2.4-mm insertion depth 

and 500-rpm rotational speed had an average lap 

shear strength of 2.6 kN, much lower than those 

made at 1000 (3.9kN), 1500 (4.1 kN) and 2000 (3.4 

kN) rpm. The results indicate that the slower 

rotational speed could not generate enough frictional 

heat and/or stirring to produce satisfactory joints. 

FY 2006 Progress Report 

Figure 1. Lap shear strength vs tool rotational speed in 

spot friction welding of AZ31 at two different insertion 

depths (2.4 and 2.8 mm) and two different surface conditions (as received and oxide layer removed). 

Figure 2. Tension-tested lap-shear specimens (top, asreceived and bottom, oxide layer removed). The number 

shown on each figure refers to the tool rpm. The insertion 

depth was 2.4 mm. 

Spot-friction-welded joints produced from 

specimens with surface oxide cleaned had higher lap 

shear strength than did as-received specimens. This 

was especially evident on the specimens with a 

93 
FY 2006 Progress Report 

2.8-mm insertion depth. The differences were 

approximately 2 kN (at 1500 rpm) to 1.5 kN (at 

2000 rpm); however, the difference was less for the 

specimens with a 2.4-mm insertion depth. 

Comparison with Resistance Spot 

Welding 

The most widely published spot joining method 

for wrought Mg is resistance spot welding (RSW). 

The study was first pioneered in 1953 by Klain et 

al.4

 and later adopted as an industrial standard with 

recommended equipment, electrodes, and process 

parameters.2,3

 The Resistance Welder Manufacturers 

Association (RWMA)2

 lists the minimum average 

tensile shear strength requirements for various Mg 

alloys and thicknesses. The data were generated 

using two different RSW equipment setups, 

including a 3-phase frequency converter machine 

and a single-phase alternating current machine 

(Figure 3). The minimum average tensile shear 

strength, using RSW for Mg specimens 1.58 mm 

thick, was between 2.4 kN (single-phase A/C 

machine) and 3.6 kN (3-phase machine with postheat). All data for spot friction welding joint strength 

from this study are overlaid and presented in Figure 

3. It is shown, without optimization in this study, 

that joint strengths as high as 4.75 kN are possible 

using the spot friction welding process. 

Figure 3. Comparison of lap shear strengths of spotfriction-welded and resistance-spot-welded AZ31. (Data 

for resistance-spot-welded joint strength are from ref. 4.) 

The lap-shear strength, t-peel strength and 

fatigue performance of spot friction welds of Al are 

as good as or better than those of RSW, self-piercing 

rivets (SPR), or clinching.5,6 In a recent cost analysis, 

High Strength Weight Reduction Materials 

costs were favorable for spot friction welding 

compared with RSW and SPR. Spot friction welding 

does not require a large power source, cooling 

water, or a compressed air supply as does RSW. The 

electrical consumption of spot friction welding is 

less than 1/20 th that of RSW.7,8

 Other merits of 

SFW include long tool life, high productivity, high 

reliability, and better working environment.7 

Conclusions 

Lap joints of 1.58-mm-thick Mg alloy 

AZ31B-O sheet were produced by spot friction 

welding. Tensile-shear testing showed that joint 

strengths of up to 4.75 kN were obtained. The spot 

friction weld strengths agree favorably with 

minimum spot weld strengths specified by the 

RWMA for similar Mg alloy sheet using the 

recommended 15.9-mm-diameter RSW. The 

removal of surface oxides from the sheets prior to 

welding increased lap shear strengths by about 50% 

at the 2.4-mm insertion depth, and it promoted 

failure by nugget pullout rather than by interface 

separation. 

Friction Stir Surface Modification of 

AM60B Magnesium Die Castings 

Experimental 

Die-cast plates with dimensions of 127×25×6 

mm were made of the AM60B by a cold chamber 

casting process under protective atmosphere. The 

nominal chemical composition of the plates was 

(5.4–6.1) Al + 0.13 Mn wt%. They contained about 

5% porosity on average from entrapped gas, which 

is typical of the HPDC process used to cast them. 

The surfaces were left in the as-cast condition rather 

than being ground or polished before FSP. 

FSP was conducted on a Bridgeport milling 

machine fitted with a 1.1-kW electric motor. The 

position of the milling machine spindle was fixed 

relative to its table so that the stir processing was 

conducted in a displacement-control mode. The stir 

tool was made of H13 steel with a shoulder diameter 

of 8 mm. The pin was cylindrical with a hemispherical tip; its dimensions were 3 mm diameter × 2 mm 

length. The working surfaces of the stir tool were 

smooth. Two tool rotation speeds were used: 1250 

and 2500 rpm. The translation speed was fixed at 1.7 

mm/s throughout the experiments. These conditions 

were used to make stir passes with lengths of about 

94 
High Strength Weight Reduction Materials 

100 mm. Some testing and analysis was done using 

single stir passes. Other plates were processed with 

5–6 passes overlapped on intervals of about 2 mm. 

Overlapping the passes created relatively consistent 

stir-processed volumes in the plates with dimensions 

of about 1.5–2×10×100 mm. 

The room temperature properties of the FSP 

surfaces were measured by Vickers microhardness 

testing and tensile testing. The microhardness 

measurements were made on metallographically 

prepared specimens taken to view the surfaces of stir 

passes. The indentations were made under a 50-g 

load in 200×200 μm arrays extending from the stir 

zones into the base metal. 

The bars with the overlapped passes were used 

to make tensile specimens. For the tensile 

specimens, blanks were cut by electrical discharge 

machining (EDM) from the stir zones using the 

shape specified in ASTM E8 for rectangular subsize 

specimens. Slices 2 mm thick were then EDM cut 

from both surfaces of the blanks to provide one 

specimen of base metal and one specimen where the 

gage section was entirely within the FSP material. 

The gage dimensions were 1 mm thick × 6.2 mm 

wide × 25 mm long. The nominal strain rate for the 

tensile tests was 1×10−3

/s. 

Results and Discussion 

A cross-sectional view showing the appearance 

of a single-pass surfacing layer made at 2500 rpm is 

shown in Figure 4(a), and a similar view of a 

multipass layer made at 1250 rpm is shown in Figure 

4(b). In both cases, some relatively large pores are 

visible in the base metal, but none were found in the 

stir zone. The base metal pores are characteristic of 

those commonly found in die-cast AM60. This 

indicates that the heat from FSP was not sufficient to 

promote the growth of pores in the heat-affected 

zone regions or the reheated stir zones of prior 

passes. The microstructure in the stir zone of the 

single-pass layer is shown in Figure 4(c). The stir 

zone consisted primarily of very fine equiaxed 

grains with sizes in the range of 5–10 μm. There is 

no evidence of solidification structure, and the 

material in the stir zone appears to contain only 

relatively small second phase particles. This grain 

FY 2006 Progress Report 

(a) 

(b) 

(c) 

Figure 4. Optical micrographs showing cross section 

views of (a) a single stir pass in AM60B made at 

2500 rpm; (b) five overlapped stir passes made at 

1250 rpm, and (c) the microstructure of the single stir 

pass. 

structure was also typical of that observed in the 

multipass layers. 

Stir zones were also examined in a scanning 

electron microscope (SEM). The micrographs 

presented in Figure 5 more clearly illustrate the 

effects stir processing had on the AM60B 

microstructure. The base metal microstructure 

[Figure 5(a)] shows the two major phases expected 

in AM60B die castings, the Mg matrix that forms 

the continuous phase plus large particles of 

Mg17Al12. 

The microstructure near the horizontally 

oriented boundary between the base metal and the 

stir zone is shown in Figure 5(b). At this location the 

shape change of the Mg17Al12 particles indicates the 

95 
FY 2006 Progress Report 

extensive deformation associated with the stir 

processing. In addition, the amount of Mg17Al12 

decreases as the boundary between base metal and 

stir zone is traversed. The microstructure in the stir 

zone is shown in Figure 5(c). At this level of 

examination, only a very small number of the large 

Mg17Al12 particles were visible in the stir zone 

microstructure, as most of this phase was apparently 

dissolved in the Mg matrix by the combination of 

heat and deformation during the stir processing. 

Limited examinations of stir zone 

microstructures were also conducted by transmission 

electron microscopy (TEM). The overall 

microstructure of a single-pass layer at this higher 

level of detail is illustrated by Figure 6(a). The TEM 

examinations confirmed that the stir zones consisted 

largely of an equiaxed Mg solid solution phase with 

grain size in the range of 5 μm. Also, numerous 

particles at a size range of several tenths of microns 

were distributed throughout the microstructure. 

These were not uniquely identified, but their number 

and relatively wide distribution indicate that they 

must be Mg17Al12. More detail of the stir zone 

microstructure is presented in Figure 6(b). At this 

location several relatively large particles are visible, 

some situated on grain boundaries. In addition, 

smaller particles can be seen distributed more 

uniformly in the matrix phase. Based on the work of 

Clark,9

 it can be concluded that the smaller, more 

uniformly distributed particles are also Mg17Al12. 

The general features of the stir zone microstructure 

regarding grain size and the disappearance of large 

Mg17Al12 particles are consistent with those made by 

Esparza et al.10 in thixomolded AM60B. 

The results from microhardness testing of a 5­

pass surfacing layer are presented in Figure 7. An 

optical micrograph is shown on the top and an image 

representation of microhardness distribution on the 

bottom. The arrays of indents used for the 

microhardness measurements are also visible in the 

micrographs and the image representations. For the 

surface of the 5-pass layer, the average hardnesses 

were 53.0 ± 7.2 kg/mm2

 in the base metal and 66.2 ± 

5.3 kg/mm2

 in the stir zone. For this case, the stir 

zone experienced an increase in hardness of 25% 

over that of the base metal. Similar measurements 

High Strength Weight Reduction Materials 

(a) 

(b) 

(c) 

Figure 5. Scanning electron micrographs showing the 

AM60B microstructure in (a) base metal, (b) near the 

boundary between the base metal (bottom) and the stir 

zone (top), and (c) the stir zone. Specimen is 

unetched. Paricles of Mg17Al12 are indicated by 

arrows in (b) and (c). Indentations from the hardness 

testing are also visible. 

96 
High Strength Weight Reduction Materials FY 2006 Progress Report 

(a) 

(b) 

Figure 6. Transmission electron micrographs in a 

single-pass surfacing layer on AM60B. General 

structure of equiaxed grains containing Mg17Al12 

particles (arrowed) is shown in (a). Details in (b) 

show some larger particles on grain boundaries as 

well as smaller particles distributed throughout 

the matrix. 

were also made on the surface of a 5-pass layer 

made with 2500 rpm. The hardness values for this 

processing condition were 56.9 ± 6.8 kg/mm2

 in base 

metal and 66.0 ± 5.7 kg/mm2

 in the stir zone. 

Summary 

The ability of FSP to produce stir passes of high 

integrity on AM60B was demonstrated. Grain 

structure in the stir zones was largely equiaxed, with 

grain sizes of 5–10 μm. Optically, the grains 

appeared to be all of a single phase. TEM 

observations revealed a homogeneous microstructure 

comprising Mg solid solution grains containing 

small second-phase particles. Based on 

metallographic evidence and previously reported 

thermodynamic considerations, the stir zone 

microstructures were concluded to consist mainly of 

the Mg matrix containing much smaller Mg17Al12 

particles than those found in the as-cast material. 

Microhardness testing indicated that stir processed 

regions experienced hardness increases up to 25% 

above base metal. 

CRADA with Ford Motor Company: 

FSW of Lightweight Vehicle Structures 

The purpose of this CRADA is to establish the 

FSW process as a viable candidate for construction 

of lightweight substructures for trucks and cars, including engine cradles, suspension sub frames, instrument panel supports, and intake manifolds. 

The approach is to conduct welding trials to establish acceptable process parameters for producing 

quality joints that maximize the mechanical performance appropriate for specific applications (e.g., 

strength, fatigue). The work will be conducted primarily on castings and extrusions of Al and Mg alloys. The project is structured in three phases: 

coupon testing is used to establish basic behaviors, 

concept testing to demonstrate commercial viability, 

and component testing to verify that the processing 

is developed sufficiently for deployment in production. 

Summary 

The major technical issues currently being addressed include shortening weld cycle times and 

maximizing weld joint tensile-shear strength. For lap 

welds of 2-mm-thick material, weld times of less 

than 5 s are desired. The target value for tensileshear strength is a minimum of 4 kN. Another major 

component of the effort relates to establishing the 

influence of tool geometry and tool surface features 

on the process adaptability and joint appearance. 

The material being used in this initial effort is 

the Al alloy 5754, which is under consideration for a 

variety of structural components. Both 2-mm and 3­

mm-thick sheets are being spot welded using several 

tool designs, including a tool designated as a reversed shoulder spiral tool (RSST), shown schematically in Figure 8. There are two unique features 

of the RSST. One is that the shoulder tapers away 

from the pin tip, opposite to the more conventional 

shape where the shoulder taper forms a shallow cav97 
FY 2006 Progress Report High Strength Weight Reduction Materials 

Figure 7. Optical micrograph showing top view of a 5-pass friction stir layer on the surface of an AM60B plate. An 

image representation of the microhardness distribution is shown on the bottom. 

Figure 8. Schematic of reversed 

shoulder spiral tool shape. 

ity. The second important feature is that the length 

of pin extension is adjustable. 

Example results from the evaluation of the influence of process (i.e., weld cycle) time and pin length 

on shear strength are shown in Figure 9. These spot 

welds were made in 2-mm-thick 5754 using insertion depths of 2.6 and 3.0 mm. Three pin lengths 

were used, 1.4, 1.2, and 1.0 mm. The overall shoulder diameter of this tool is 10 mm. These results indicate that at the shorter processing times, the 

intermediate pin length of 1.2 mm produced the 

strongest welds. For the 1.2-mm pin, the effect of 

insertion depth was insignificant. Although additional testing is required to build statistical confidence in the results, the joint strengths were 

relatively insensitive to processing time. This suggests that increasing weld cycle time may not be 

necessary for meeting minimum strength requireFigure 9. Variation of lap-shear strength with process 

(welding) time for joints of 2-mm-thick 5754 spot welded 

with an insertion depth of 2.6 mm and pin lengths of 1.0 

and 1.4 mm. 

ments, a finding somewhat contrary to the expected 

result. 

Figure 10 shows a compilation of results on the 

effects of insertion depth and pin length on lap-shear 

strength. All of these spot weld joints were made 

with a 5 s weld time. Figure 10 indicates that for 

each pin length from 0-2 mm, a peak occurred in the 

variation of lap-shear strength with insertion depth. 

However, these strength peaks occurred over narrow 

ranges of insertion depth. It appears that it is feasible 

to obtain a joint strength that exceeds the desired 

minimum of 4 kN. The narrow range of insertion 

depths over which the peaks occur means 

98 
High Strength Weight Reduction Materials FY 2006 Progress Report 

Figure 10. Variation of lap-shear strength with insertion 

depth for a weld time of 5 s and pin lengths of 0-2 mm. 

that the process must be very closely controlled. 

This is undesirable from a production perspective, 

and efforts to increase strength while broadening the 

strength peak are continuing. 

The strength data presented in Figure 9 are also 

replotted in Figure 11, where the reaction forces on 

the stir tooling are included. This plot shows that the 

tool reaction forces decreased as processing times 

increased from 1–5 s. The effect of processing time 

on the reaction force is significant because it will be 

a factor in determining the sizes of robots needed for 

producing welds using portable welding heads. The 

13-18 kN reaction forces measured at lower 

processing times are quite high compared with the 

levels used for RSW robots. The implications of 

these observations are being discussed with 

manufacturers of robotic equipment. 

Figure 11. Variations of both lap-shear strength and stir 

tool reaction forces with process (welding) time for spot 

welds of 2-mm-thick 5754. 

Conclusions 

Progress has been made in FY 2006 in FSW/P 

of lightweight materials for automotive structure 

applications. It has been demonstrated that, for both 

Mg and Al alloys, friction stir spot welding was capable of producing welds exceeding the strength 

requirements for RSW of the same material and 

thickness. The influence of tool geometry and process conditions on the weld joint strength of Al5754 

was investigated, and the relatively high process 

load to meet the strength requirement needs to be 

considered in the design and fabrication of welding 

robots for assembly line application of the friction 

stir spot welding process. 

The ability of FSP to produce a high-integrity 

surface layer on AM60B was demonstrated. In addition to removing the solidification porosity of the 

cast AM60B, FSP regions experienced hardness/strength increases of up to 25% above base 

metal, suggesting the possibility of considerable improvement of mechanical properties (strength, ductility and fatigue life) in the FSP region. 

Presentations/Publications 

Z. Feng, M. L. Santella, S. A. David, R. J. Steel, 

S. M. Packer, T-Y Pan, M. Kuo, and R. S. Bhatnagar, “Friction Stir Spot Welding of Advanced HighStrength Steels—A Feasibility Study,” SAE 2005 

Congress, SAE-International, Detroit, 2005-2001­

1248, 2005. 

M. Awang, V. Mucino, Z. Feng, and S. A. 

David, “Thermo-Mechanical Modeling of Friction 

Stir Spot Welding (FSSW) Process: Use of an Explicit Adaptive Meshing Scheme,” SAE 2005 Congress, SAE-International, Detroit, 2005-2001-1251, 

2005. 

S. G. Arul, T-Y Pan, J. Pan, P. C. Lin, Z. Feng, 

and M. L. Santella, “Microstructures and Failure 

Mechanisms of Spot Friction Welds in Lap-Shear 

Specimens of Aluminum 5754 Sheets,” SAE 2005 

Congress, Detroit, 2005-2001-1256, 2005. 

M. L. Santella, Z. Feng, C. Degen, and T-Y Pan, 

“Modifying AM60B Magnesium Alloy Die Cast 

Surfaces by Friction Stir Processing,” in 7th International Conference on Trends in Welding Research, 

S. A. David. Ed., TMS, Pine Mountain, GA, 2005. 

M. L. Santella, Z. Feng, C. Degen, and T-Y Pan, 

“Friction Stir Processing to Improve the Surface 

Properties of Aluminum and Magnesium Castings,” 

99 
FY 2006 Progress Report High Strength Weight Reduction Materials 

Materials Science & Technology 2005 Conference, 

ASM International, Pittsburgh, 2005. SAE Technical Paper 982362, Society of AutoM. L. Santella, T. Engstrom, D. Storjohann, and motive Engineers, Warrendale, PA, 1998. 

T-Y Pan, “Effects of Friction Stir Processing on Me­ 6 M. Fujimoto, M. Inuzuka, M. Nihio, Y. Nakachanical Properties of the Cast Aluminum Alloys shima, “Development of Friction Spot Joining 

A319 and A356,” Scripta Materialia 53(2), 201–206 (Report 2): Mechanical Properties of Friction 

(2005). Spot Joints,” Preprints of the National Meeting 

Z. Feng, T-Y Pan, M. L. Santella, and S. A. of Japan Welding Society, no.74, pp.6–7, 2004. 

David, “Friction Stir Spot Welding and Its Applica­ 7 “Mazda Develops World’s First Aluminum 

tions,” AeroMat 2005, ASM International, Orlando, Joining Technology Using Friction Heat,” 

FL, 2005. Mazda Media Release, February 27, 2003. 

W. Woo, P. K. Liaw, H. Choo, D. W. Brown, Z. 8 R. Hancock, “Friction Welding of Aluminum 

Feng, S. A. David, and C. R. Hubbard, “D-41 Two- Cuts Energy Cost by 99%,” Welding Journal 83, 

Dimensional Mapping of Residual Stresses in 6061­ 40 (2004). 

T6 Aluminum Friction Stir Welds,” Powder Diffrac- 9 J. B. Clark, “Age Hardening in a Mg-9 wt% Al 

tion 19(2), 202 (2004). Alloy,” Acta Metallurgica 16, 141–152 (1968). 

M. L. Santella, A. Frederick, C. Degen, and T-Y 10 J. A. Esparza, W. C. Davis, and L. E. Murr, 

Pan, “The Use of Friction-Stir Technology to Mod­ “Microstructure-property Studies in Frictionify the Surfaces of AM60B Magnesium Die Cast- stir-welded, Thixomolded Magnesium Alloy 

ings,” JOM 58(5), 56–61 (2006) AM60, Journal of Materials Science 38, 941– 

M. L. Santella, T-Y Pan, and C. Degen, “Modi­ 952 (2003). 

fying AM60 Magnesium Alloy Die Cast Surfaces by 

Friction Stir Processing,” presented at SAE 2006 

World Congress, Detroit, April 6, 2006. 

T-Y Pan, M. L. Santella, P. K. Mallick, A. Frederick, and W. J. Schwartz, “A Feasibility Study on 

Spot Friction Welding of Magnesium Alloy AZ31,” 

in Proceedings of 2006 World Magnesium Conference, 2006. 

References 

1 A. Luo, “Wrought Magnesium Alloys and 

Manufacturing Processes for Automotive Applications,” SAE Technical Paper 2005-01-0734, 

Society of Automotive Engineers, Warrendale, 

PA, 2005. 

2 Resistance Welding Manual, revised 4th ed., 

Resistance Welder Manufacturers' Association, 

2003. 

3 Metals and Their Weldability, pp. 418–424 in 

Welding Handbook, 7th ed., vol. 4, American 

Welding Society, Miami, 1982. 

4 P. Klain, D. L. Knight, and J.P. Thorne, “Spot 

Welding of Magnesium with Three-Phase Low 

Frequency Equipment,” Welding Journal 32(1), 

7–18 (1953). 

5 S. Kallee and D. Nicholas, “Application of Friction Stir Welding to Lightweight Vehicles,” 

100 
High Strength Weight Reduction Materials FY 2006 Progress Report 

F. Basic Studies of Ultrasonic Welding for Advanced Transportation Systems 

Principal Investigator: Zhili Feng 

Oak Ridge National Laboratory 

1 Bethel Valley Road, Oak Ridge, TN 37831 

(865) 576-3797; fax: (865) 574-4928; e-mail: fengz@ornl.gov 

Chief Scientist: James J. Eberhardt 

(202) 586-9837; fax: (202) 587-2476; e-mail: james.eberhardt@ee.doe.gov 

Field Technical Manager: Philip S. Sklad 

(865) 574-5069; fax:(865) 574-6098; e-mail: skladps@ornl.gov 

Participants: 

X. Zhang and M. L. Santella, Oak Ridge National Laboratory 

W. Zhang and K. Li, Oakland University 

Contractor: Oak Ridge National Laboratory 

Contract No.: DE-AC05-00OR22725 

Objective 

• Develop a fundamental understanding of the ultrasonic welding (UW) process. 

• Establish process conditions to optimize joint properties of aluminum (Al) alloys and other lightweight materials for auto body applications. 

• Explore opportunities for joining dissimilar materials. 

• Explore the feasibility of ultrasonic processing in other novel and unique situations in materials processing, 

such as consolidating metal powder, producing functionally graded components, and modifying surface properties. 

Approach 

• Perform UW experiments to develop correlations of process parameters with joint properties. 

• Characterize details of joint microstructures. 

• Model the fundamental interactions of ultrasonic waves with solids. 

Accomplishments 

• Demonstrated UW of magnesium (Mg) alloys. 

• Developed an efficient computational model to study the acoustic energy distribution in the welded workpiece. 

• Investigated the geometric effects on ultrasonic energy generation and dissipation during the welding process. 

• Initiated experimental study on residual stress distribution in the bonding region. 

Future Direction 

• Continue welding process development for dissimilar materials. 

• Study the structure-property relationship of joints. 

101 
FY 2006 Progress Report 

Introduction 

Ultrasonic welding (UW) uses high-frequency 

mechanical vibrations to produce a solid-state metallurgical bond (weld) between metals. The basic 

process setup is illustrated in Figure 1 for typical 

spot weld configurations. An electro-mechanical 

converter converts high-frequency electric current to 

mechanical vibration. The mechanical vibration is 

then modulated and amplified by the booster/horn 

before it is applied to the overlapping workpiece 

through the vibration of the sonotrode. A moderate 

clamping force is applied to ensure the mechanical 

vibration is transferred to the sheet-to-sheet interface 

(the faying surface) where the weld is created. Typically, the mechanical vibration is at 20–40 kHz with 

an amplitude range of 5–50 μm. The power delivered to the workpiece is in the range of several hundred to several thousand watts. 

Figure 1. Schematic setup of UW process. 

Figure 2 illustrates the essential underlying 

mechanisms involved in the formation of the metallurgical bond at the workpiece interface during UW. 

The combination of normal clamp pressure and lateral mechanical vibration results in several important 

effects at the workpiece interface. The lateral mechanical vibration introduces relative motion at the 

interface. The frictional action at the interface due to 

the relative motion and the normal pressure breaks 

the surface oxides and other contaminants. As a result, clean metal surfaces are brought into contact 

with each other under pressure. Frictional 

High Strength Weight Reduction Materials 

Figure 2. Schematic representation of bond interface before (left) and after (right) UW. 

heating also occurs at the bond interface. The heating promotes both localized deformation and diffusion in the region where the pressure is applied. 

When the process conditions are right, a metallurgical bond is formed without melting at the bonding 

interface. The self-cleaning nature of UW and its 

ability to form metallurgical bonds without melting 

are two important advantages of the process. 

UW creates a joint without bulk melting. Thus, 

UW is inherently immune to welding defects associated with the solidification in various fusion welding 

processes (electric resistance spot welding, arc welding, laser welding, etc.). Solidification-induced defects such as solidification cracking and porosity can 

be a major quality problem for a number of highperformance lightweight structural materials used in 

automotive body structures. In addition, UW offers 

the possibility for joining dissimilar metals such as 

Al to steel and Al to Mg, metal-to-metal composites 

that have been extremely difficult for conventional 

fusion welding processes. Because melting is 

avoided, the UW process is much more energy efficient. UW has attracted considerable attention in the 

auto industry as an important enabling joining technology for assembling structural components made 

of Al alloys, Mg, and other high-performance lightweight materials. 

Progress before FY 2006 

UW of Al alloys was successfully demonstrated 

in FY 2003, the first year of the program. UW of 

other materials was also explored. Studies on dissimilar welds between Al2024 and Al6061 revealed 

a number of important phenomena at the interface 

under different process conditions. 

The FY 2004 research involved studying the relationship between the weld quality of Al alloys and 

weld process conditions. It was observed that the ultrasonic weld quality strongly depends on a number 

of geometric factors, such as the weld location, the 

102 
High Strength Weight Reduction Materials 

shape and size of the workpiece, and the sonotrode 

vibration direction relative to the workpiece. Under 

certain circumstances, cracking occurs at unexpected 

locations. Destruction of prior welds by subsequent 

welding was also observed. These observations suggested that the acoustic wave propagation and reflection in the workpiece being welded are very complicated and would result in considerable weld-to-weld 

quality variations. 

In FY 2005, the program focused on investigating the acoustic energy distribution during welding. 

By utilizing a real-time infrared (IR) thermography 

technique, localized temperature rise in the workpiece during UW was observed and recorded as a 

function of time. The influence of the geometry factors on the energy distribution was experimentally 

investigated to elucidate possible causes for the observed weld quality variations. The causes for the 

formation of unforeseen hot spots and cracking were 

identified, and the factors contributing to the acoustic energy generation and dissipation in welded samples determined. In addition, we began to formulate 

an efficient computational modeling approach to further the understanding of the acoustic energy distribution in the workpiece. This model makes it possible to investigate the localized acoustic energy distribution and heat generation during ultrasonic welding process. 

The research activities in FY 2006 included investigating the feasibility of UW of Mg alloys, because Mg alloys used for auto body structure material received increased attention in the automotive 

industry. We also completed the development of the 

computational model for acoustic energy distribution 

in the workpiece during UW. In addition, the residual stresses in the UW joint were studied using an 

interferometric strain/slope rosette (ISSR) stress 

measurement technique. 

Joining of Mg Alloys 

Two Mg alloys (AZ31B and AM50) were studied. AZ31B rolled sheet had a nominal thickness of 

1 mm, and the die-cast AM50 sheet was 2 mm thick. 

The weld coupons were 25 mm wide. A standard 25-

mm overlap between the two workpieces was used 

in the welding trials. 

All welding tests were performed on a Sonobond 

ultrasonic welder. The machine operated at 20 kHz 

and had maximum rated power output of 2500 W. A 

series of welding tests were performed. The welding 

FY 2006 Progress Report 

time varied from 0.7 to 1.75 s, and the welding energy varied from 980 to 2400 J. 

In addition to same-alloy welding (AZ31B to 

AZ31B and AM50 to AM50), welding between 

AZ31B and AM50 was also performed. Figure 3 

shows the appearance of typical welds made in this 

study. 

Figure 3. Appearance of ultrasonic welds. Top weld: 

AZ31B to AZ31B (t=1 mm); middle: 2 mm AM50 to 1 

mm AZ31B; bottom: AM50 to AM50 (t-2 mm). 

Excellent metallurgical bonding was obtained 

for both Mg alloys. Figure 4 shows the defect-free 

4-mm-wide bonding interface of AM50. At higher 

magnification (Figure 5), severely elongated grains 

are formed along the bonding interface because of 

the extensive deformation of interface during UW. 

The width of the deformed grain zone is very narrow, approximately 20 μm, indicating UW involving 

very localized deformation at the interface. This was 

consistent with UW of other materials such as Al alloys studied in the past. 

Outside the bonding region, randomly distributed solidification porosity/voids are clearly visible 

in the base metal of AM50, as shown in Figures 4 

and 5. Such solidification porosity/voids are common and associated with the die-casting process. 

Figure 6 shows the bonding interface of the dissimilar weld between AZ31B and AM50. It is evident that metallurgical bonding between the two different Mg alloys was produced by the UW process. 

While we did not perform mechanical testing to 

determine the joint strength (planned in FY 2007), 

the peel test commonly used to test resistance spot 

103 
FY 2006 Progress Report High Strength Weight Reduction Materials 

Figure 4. Weld cross-section of AM50. 

Figure 5. Microstructure at the bonding interface of 

AM50. Note the solidification porosity in the base metal 

region. 

welds was conducted to determine the quality of ultrasonic weld. As shown in Figure 7, the ultrasonic 

welded joint failed in the base metal adjacent to the 

weld, indicating that the ultrasonic weld had adequate strength. 

Modeling UW by Means of Time Domain 

Elastic Wave Solution Technique 

We have completed the development of a UW 

model based on the time-domain elastic wave solution technique. The theoretical basis and modeling 

results are summarized here. 

Wave propagation and the vibration modes in 

elastic plates and cylinders have been model problems for solutions of elastic wave equations. Early 

Figure 6. Microstructure at the bonding interface between AZ31B (top) and AM50 (bottom). 

Figure 7. Base metal failure during peel test AZ31B 

ultrasonic weld. 

works by Lord Rayleigh1,2

 considered cases where 

the thickness H of a plate is much smaller than the 

dimensions in the other two directions; and the extensional modes, which involve wave propagation 

along the thickness of the plate, can be neglected. In 

the more general case of thin plates with finite 

thickness, a series expansion in the thickness of the 

plate can be obtained, which can be solved numerically.

3,4

 These early studies focused on free vibrations. 

For thin plates driven by external periodic 

forces, the solutions are often expressed in terms of 

a summation over different frequency modes.5,6

 Research of driven plates so far has mostly focused on 

the steady-state problems. Little attention has been 

given to non-steady-state cases. 

The short duration of the UW process, a fraction 

of a second, contains only tens of thousands of cy104 
High Strength Weight Reduction Materials FY 2006 Progress Report 

cles at the driving frequency. This is not long 

enough to establish steady-state vibrations in the 

plates. A detailed study of a driven elastic plate in 

the time domain is necessary. 

Although the welding process itself is not an 

elastic process and cannot be described by linear 

elastic equations, the distribution of the acoustic energy over the driven plate can be described reasonably well by the linear equations, if we approximate 

the energy dissipation by a damping parameter over 

the entire plate. For the simplicity of the model, we 

also neglect the effect of heating and the transfer of 

energy from the top plate to the bottom plate. Again, 

this energy can also be modeled by an appropriately 

chosen damping term within the linear theory. These 

approximations yield a time-dependent linear acoustic problem for a single driven plate that can be readily solved using a finite-element numerical technique. 

Governing Equations of Elastic Vibration of 

Thin Plates pertaining to UW 

Consider a thin plate situated in a threedimensional (3-D) rectangular coordinate system, (x, 

y, z), with the thickness in z direction. The plate is 

bounded by the planes z = h and z = ! h and the 

thickness of the plate is H = 2h at the start. Let U, V, 

W be the displacements of the plate in x, y, z directions respectively. We followed the elastic modeling 

approach for high-frequency extensional vibrations 

of plate by Kane and Mindlin.3

 It was assumed that 

the displacements in x and y directions are independent of the thickness of the plate and the displacement in z direction varies linearly along z with 

no displacement at the middle plane of the plate. 

Hence, the three displacements follow the formulas 

U (x, y, z,t) = u(x, y,t)

V (x, y, z,t) = v(x, y,t) (1) 

W (x, y, z,t) = z w(x, y,t) h 

In Eq. (1), the equation of motion can be solved 

in the reduced domain (x, y, t). The corresponding 

(λ + 2G)uxx + Guyy + (λ + G)vxy + π λw + C1e 

−α[(x−xs 

x 

plate-displacement system of equations of motion is 

(λ + 2G)uxx + Guyy + (λ + G)vxy + π λwx +

T1 = ρutt 3H H

Gvxx + (λ + 2G)vyy + (λ + G)uxy + π λwy +

T2 = ρvtt 3H H

G(wxx + wyy ) − 3π λ(ux + vy ) − H

π2

2 (λ + 2G)w +

3T3 = ρwtt H H 

(2) 

where λ, G are Lame’s constants; ρ is the mass density; H = 2h is the thickness of the plate; T1 ,T2 ,T3 

are the components of surface traction in x, y and z 

directions. 

We represent the driving vibration of the 

sonotrode by a simple temporal harmonic oscillation 

of surface traction with exponentially decaying amplitude in the spatial domain. The choice of a continuous surface traction is to avoid difficulties in 

solving Eq. (2) arising from spatial discontinuities. 

In an UW process, strong friction exists between the 

workpiece interface. A simple friction term is used 

in each equation of motion to represent the friction 

effect. This leads to Eq. (3), the elastic model for 

UW. 

In Eq. (3), ω is the vibration angular frequency 

of the force applied to the weldment, (xc , yc ) is the 

center of the sonotrode on the workpiece, and 

C1 ,C2 ,C3 ,α,γ are nonnegative constants. 

There are two physical origins of the damping 

coefficient γ in Eq. (3). The first is the obvious 

source, the dissipation of the elastic energy into heat 

on the plates. The second is less obvious but equally 

important. This is the energy leak due to imperfect 

impedance match between the sonotrode, the workpiece, and the supporting anvil. 

Boundary Conditions 

We consider a rectangular plate [0, A]×[0, B], 

corresponding to the upper workpiece in the UW 

experiment. In the UW process, the sides of weldment are free. Hence, at the four sides of the rectangular plate, we propose zero normal stresses. Since 

the plate is thin, we propose zero shear stresses 

along the z-direction (the thickness direction). Since 

)

2 +(y−ys )2 ]

sinωt = ρutt + γut 3H 

Gvxx + (λ + 2G)vyy + (λ + G)uxy + π λwy + C2e 

−α[(x−xs )

2 +(y−ys )2 ]

sinωt = ρvtt + γvt 

(3) 

3H 

G(wxx + wyy ) − 3π λ(ux + vy ) − π2(λ + 2G)w + C3e 

−α[(x−xs )

2 +(y−ys )2 ]

sinωt = ρwtt + γwt H H 2 

105 
FY 2006 Progress Report 

the plate is clamped along the z-direction at the 

sonotrode and the vibration in the surface traction is 

in the xy-plane, the rotation movement of the displacement vector <U, V, W> about the z-axis is negligible at the boundary. Hence, we propose the irrotational condition of the displacement vector <U, V, 

W> about the z-axis. The resulting boundary conditions for Eq. (3) are 

(λ + 2G)ux + λvy +

πλ w = 0 

h 12 

u − v = 0 (4) y x 

w = 0 x 

at the boundaries x = 0 and x = A, and 

λux + (λ + 2G)vy +

πλ w = 0 

h 12 (5) u − v = 0 y x 

w = 0 y 

at the boundaries y = 0 and y = B. The compatible 

boundary conditions at a corner of the rectangular 

domain require 

πλ u x = v y = −

h(λ + G)2 12 w 

(6) u y − v x = 0

w x = w y = 0

The initial conditions for Eq. (3) in our UW 

model 

u(x, y,0) = v(x, y,0) = w(x, y,0) 

= ut (x, y,0) = vt (x, y,0) = wt (x, y,0) (7) 

= 0

Eqs. (3)–(7) form a closed mathematical system 

of an initial boundary problem of a system of linear 

partial differential equations. 

The direction of the applied vibration in UW is 

controlled by the parameters C1 ,C2 ,C3 in Eq. (3). For 

vibration along the length of the plate (in x direction), we assign C1 = 0.035,C2 = C3 = 0. For vibration 

along the width of the plate (in y direction), we assign C1 = 0,C2 = 0.035,C3 = 0 . The coefficient of the 

decay of the applied vibration in Eq. (3) is set as α = 

32000 so that the applied force vanishes quickly 

away from the welding spot centered at (xc , yc ). 

Finally, the following material properties were 

used for the Al alloy workpiece. 

High Strength Weight Reduction Materials 

E = 7.0404172 N 

2 ,σ = 0.34,ρ = 2700 kg 3 , m m

λ = σE N 

2 ,G = E N 2 , (1+ σ)(1− 2σ) m 2 1( + σ) m

H = 0.001m, A = 0.1m,B = 0.025m, 

xs = 0.065m, ys = 0.0125m,ω =169646.01Hz 

Numerical Simulations and Comparisons to 

the Experimental Results 

A variable-coefficient ODE solver, 

DVODE_F907,8

 was used in the computation. The 

stiff backward differentiation formulas were adopted 

as the integration method. 

Consistent with the UW experiments conducted 

in FY 2005, the Al alloy workpiece was 0.025 m 

wide and 0.1 mm long. The sonotrode was located at 

xc=0.075 m and yc=0.0125 m. 

We assume that the temperature distribution 

during UW as observed in IR measurement is in 

proportion to the sum of the kinetic and potential 

energy distributions on the plate. With an applied 

periodic vibration force, the solution and energies 

change approximately periodically. Hence, we define a time averaged energy sum as 

E(x, y) = T 

1 T 

∫

0 

(K (x, y,t) + P(x, y,t))dt = EK + EP 

(8) 

where K(x, y,t), P(x, y,t) are the kinetic and potential 

energies and T is a sufficiently long time. In our 

simulations, it was found that the potential energy 

dominates the kinetic energy. Hence, the distribution 

of the averaged energy sum, E, is much closer or 

similar to that of the time averaged potential energy, EP . 

First, the UW model was applied to the case in 

which the direction of the driving vibration from the 

sonotrode was along the width of the rectangular 

plate (in y-direction), with damping coefficient γ = 

20ρ/s. Figure 8 shows the out-of-plane motion of 

106 
High Strength Weight Reduction Materials FY 2006 Progress Report 

Figure 8. Snapshots at different time instances of the 

plate vibration in the thickness direction (z-axis) during 

UW. The driving vibration is in the direction of the yaxis. Damping coefficient γ = 20ρ/s. Acoustic excitation 

locates at x = 0.065 m and y = 0.0125 m. 

the plate in one cycle. The oscillations of the surface 

were mainly confined at two spots near the long 

edge of the plate, away from the applied vibration 

location (x = 0.065m, y = 0.0125m). 

Figure 9 shows the distribution of the timeaveraged energy. We observe that two high-energy 

spots appeared on the long edges of the plate, contributed largely by the potential energy. The locations of the high-energy spots qualitatively agree to 

the IR thermography image reported in FY 2005. 

Since the potential energy is high at those spots, 

cracks are likely to develop there. 

Second, the same vibration force was applied 

along the length of the rectangular plate (in xdirection), with the same damping coefficient γ = 

20ρ/s. Figure 10 shows the distribution of the timeaveraged energy. A completely different energy distribution was observed. There is only one highenergy spot located at the mid-width of the plate. 

Furthermore, the magnitude of the high-energy spot 

is much smaller than that shown in Figure 9. Such 

energy distribution is consistent with the IR thermograph observation where no hot spots were observed 

and no crack developed under this testing condition. 

Figure 9. Distribution of the time-averaged sum of the 

kinetic and potential energy density in the plate during 

UW when the driving vibration is in the direction of the 

y-axis. Damping coefficient γ = 20ρ/s. 

Figure 10. Distribution of the sum of the kinetic and potential energies on the plate during UW when the driving 

vibration is in the direction of the x-axis. Z represents the 

amount of the energy sum. Damping coefficient γ = 

20ρ/s. 

Third, the model was used to study the effect of 

acoustic impedance by changing the damping coefficient. Small damping means that most of the energy is trapped within the plate, while large damping 

means that a large portion of the energy is leaked to 

the anvil. Therefore, the damping parameters can be 

used to simulate the effect of acoustic impedances of 

the workpiece. In particular, two plates with a frictional interface correspond to a relatively small 

107 
FY 2006 Progress Report 

damping, whereas a single plate corresponds to a 

large damping. 

For the impedance study, the vibration force was 

applied along the width of the plate (in y-direction). 

Figure 11 shows the distributions of the timeaveraged energy when using a smaller damping coefficient, γ = 2ρ/s. This corresponds to larger interface impedance. The high-energy spots appeared at 

the corners of the plate away from the welding spot. 

There is no high-energy spot around or close to the 

welding spot. 

Figure 11. Distribution of the sum of the kinetic and potential energies on the plate during UW when the driving 

vibration is in the direction of the y-axis. Damping coefficient γ = 2ρ/s. 

Figures 12–14 show the cases of larger damping 

coefficients (γ = 80ρ/s, 160ρ/s and 320ρ/s). As the 

damping coefficient increases, more elastic energy 

leaked to the anvil. The effect of the plate geometry 

plays less and less of a role since the motion is 

quickly damped out away from the welding spot. In 

Figure 12, as the damping increases to γ = 80ρ/s, 

four main high-energy spots occurred at the long 

edges of the plate surrounding the welding spot as a 

center. There are also four small secondary highenergy spots closer to and on the paths of the welding spot. As the damping coefficient increases to γ = 

160ρ/s (Figure 13), the four high-energy spots on the 

long edges of the plate become secondary highenergy spots, while two main high-energy areas are 

located inside the plate and on the two sides of the 

welding spot along the direction of the applied vibration. As the damping coefficient increases to γ = 

High Strength Weight Reduction Materials 

Figure 12. Distribution of the sum of the kinetic 

and potential energies on the plate during UW 

when the driving vibration is in the direction of the 

y-axis. Damping coefficient γ = 80ρ/s. 

Figure 13. Distribution of the sum of the kinetic and potential energies on the plate during UW when the driving 

vibration is in the direction of the y-axis. Damping coefficient γ = 160ρ/s. 

320ρ/s (Figure 14), high-energy spots disappear 

from the edges of the plate and from the two highenergy spots formed on the side of the welding spot 

along the direction of vibration. The last figure is 

very similar to the experimental image of a single 

plate. 

Conclusions 

1. Two Mg alloys, AM50 and AZ31B, have been 

successfully welded using the ultrasonic welding process. Metallurgical bonding was produced. Peel tests revealed that the weld joint 

108 
High Strength Weight Reduction Materials 

Figure 14. Distribution of the sum of the kinetic and potential energies on the plate during UW when the driving 

vibration is in the direction of the y-axis. Damping coefficient γ = 320ρ/s. 

failed in the base metal region adjacent to the 

spot weld, demonstrating sufficient strength of 

the ultrasonic welds produced in this study. 

2. An efficient computational modeling approach 

based on the time domain elastic wave propagation principle has been developed to further the 

understanding of 

3. The acoustic energy distribution in the workpiece. This model makes it possible to investigate the localized acoustic energy distribution 

and heat generation during UW process. 

FY 2006 Progress Report 

References 

1 S. H. Crandall, “Rayleighs Influence on Engineering Vibration Theory,” Journal of the 

Acoustical Society of America 98(3), 1269–72, 

1995. 

2 A. E. H. Love, A Treatise on the Mathematical 

Theory of Elasticity, New York, Dover, 1944. 

3 T. R. Kane and R. D. Mindlin, “HighFrequency Extensional Vibrations of Plates,” J. 

Appl. Mech. 23, Trans. ASME 78, A-277–283, 

1956. 

4 S. S. H. Chen and T. M. Liu, “Extensional Vibration of Thin Plates of Various Shapes,” 

Journal of the Acoustical Society of America 

58(4), 828–831, 1975. 

5 C. T. Sun, and J. M. Whitney, “Forced Vibrations of Laminated Composite Plates in Cylindrical Bending,” Journal of the Acoustical Society of America 55(5), 1003–8, 1974. 

6 N. H. Farag and J. Pan, “Free and Forced Inplane Vibration of Rectangular Plates,” Journal 

of the Acoustical Society of America 103(1) 

408–13, 1998. 

7 http://www.radford.edu/~thompson/ 

vodef90web. 

8 P. N. Brown, G. D. Byrne, and A. C. Hindmarsh, “VODE: A Variable Coefficient ODE 

Solver,” SIAM J. Sci. Stat. Compu. 10, 1038– 

51, 1989. 

109 
FY 2006 Progress Report High Strength Weight Reduction Materials 

110 
High Strength Weight Reduction Materials FY 2006 Progress Report 

5. LIGHTWEIGHT VEHICLE STRUCTURES 

A. Application of Superplastically Formed Aluminum for Truck Body Panels 

Principal Investigator: Nirmal M. Tolani 

International Truck and Engine Company 

2911 Meyer Road, Fort Wayne, IN 46803 

(260) 461-1238; fax: (260) 428-3661; e-mail: nirmal.tolani@nav-international.com 

Chief Scientist: James J. Eberhardt 

(202) 586-9837; fax: (202) 587-2476; e-mail: james.eberhardt@ee.doe.gov 

Field Technical Manager: Philip S. Sklad 

(865) 574-5069; fax: (865) 574-6098; e-mail: skladps@ornl.gov 

Contractor: International Truck and Engine Company 

Contract No.: 4000026691 

Objective 

• Investigate applications of superplastic aluminum for low to moderate volume (up to 30000/year) body panels 

to provide a lightweight and low-tooling-cost alternative to steel and sheet molding compound. 

Approach 

• Select a large exterior truck body panel having a complex shape and moderate production volumes. 

• Develop part design, conduct forming simulation and finite element analysis (FEA). 

• Build prototype parts to demonstrate process capability and develop realistic part cost data. 

• Evaluate the performance of the parts in the real vehicle environment. 

Accomplishments 

The program was initiated in March 2005. To date, the following tasks have been completed. 

• Selected appropriate exterior truck body part. 

• Completed part design and developed the part support structure design concepts. 

• Completed forming simulations to verify manufacturability 

• Completed FEA of part design and support structure. 

• Completed second-iteration FEA of part support structure with the exception of joint between mounting 

structure and superplastic forming (SPF) bumper panels. 

• Completed prototype SPF tooling. 

• Completed prototype forming trials and development. 

• Completed trial of a lower-cost SPF material from a domestic supplier to reduce cost. 

• Made six prototype parts for preliminary assembly and test trials. 

111
FY 2006 Progress Report High Strength Weight Reduction Materials 

• Reassessed and revised cost data for production parts 

• Ran a preliminary cab shaker test to assess the interface between SPF panel and mounting structure. 

Future Direction 

• Productionize the prototype tool and run the part in the production process to gain a better understanding of the 

inconsistencies noted in the prototype forming trials. 

• Evaluate surface quality and material thickness variations of the product’s representative parts. 

• Get a true assessment of production cycle times and costs under the production environment. 

• Develop a robust interface between the SPF panel and mounting structure. 

• Perform cab shaker test and/or vehicle test to assess part durability for possible implementation in production. 

Introduction 

Traditionally, commercial truck (Class 5–8) cabs 

have been made from two materials—steel and 

aluminum. To stay profitable, the truck industry has 

been going through product rationalization. To 

reduce the tooling cost associated with two different 

cabs, there appears to be a trend to standardize to a 

single higher-volume steel cab. This trend will make 

the Class 7–8 vehicles cabs heavier and reduce the 

fuel efficiency. 

This trend toward conversion from aluminum to 

steel in heavy-duty trucks can be reversed if the 

formability of aluminum for producing future 

aerodynamic shapes is improved and the tooling cost 

to produce aluminum cab panels can be significantly 

reduced. Deployment of SPF technology offers the 

opportunity to do so. 

SPF of certain aluminum alloys offers the ability 

to produce complex aerodynamic shapes for truck 

body applications using lower-cost tooling. This 

technology allows the production of lightweight, 

highly integrated, net-shape components that often 

consolidate many parts in one. It reduces the number 

of parts, fasteners, and assembly operations required 

for complex truck body parts and enables the use of 

aluminum in place of steel at competitive costs. 

A typical SPF process is illustrated in Figure 1. 

The process uses a single-sided die rather than a 

matched two-sided die. The sheet blank is clamped 

in the die and blow-formed. Complex-shaped parts 

can be formed that are not otherwise possible. 

International Truck and Engine Company 

studied the feasibility of using the SPF process for 

Figure 1. Illustration of typical SPF process. 

large truck panels in 1985. At that time, the SPF 

aluminum process cycle times and superplastic 

aluminum material costs were high; also, the 

Class A surface finish required for exterior body 

panels could not be achieved. Although the tooling 

costs were low, the part costs were very high. It was 

determined that SPF was economically viable only 

for very low volumes (<1000 pieces). 

In recent years, significant new developments in 

SPF materials and processes have been reported. It is 

reported that these developments have made higher 

production volumes feasible, matching those of 

Class 7–8 truck volumes, through a combination of 

increased forming rates and lower material cost. It is 

also reported that for moderate levels of forming, the 

Class A surface condition required for exterior body 

panels is achievable. 

These recent developments make SPF 

technology increasingly attractive for heavy-vehicle 

applications. However, there is a strong need to 

validate the claims made about the forming cycle 

112
High Strength Weight Reduction Materials FY 2006 Progress Report 

times, part costs, tooling costs, Class A surface 

capability, and fatigue properties as they apply to 

large cab panels and structural parts. The only way 

to convince the heavy truck industry to deploy this 

relatively new technology is to demonstrate its 

feasibility for a large exterior body part. 

Project Goal 

The project goal is to initially demonstrate the 

feasibility of SPF technology for a large exterior 

body panel as phase 1. If the results look promising, 

in the second phase, application of SPF will be 

expanded to making more complex body subassemblies and/or the entire truck cab, depending on 

the funding level provided. 

Project Plan 

Phase 1 of the project involves the design, 

analysis, prototyping, and testing of the large 

exterior truck body part. The following task 

descriptions were proposed to address the key issues 

related to the use of SPF in heavy-duty vehicle 

applications. 

Design and analyze parts: This task has helped 

to understand design and tooling considerations and 

limitations. It helped develop guidelines for future 

designs with SPF aluminum. 

Build prototype parts: The production of 

prototype parts will help evaluate tooling issues, 

forming process cycle times, part wall thickness 

variations, and the capability of SPF aluminum to 

provide a Class A surface. 

Perform durability tests: This will help 

evaluate the assembly, stiffness and durability of 

SPF aluminum in Class 8 truck applications. 

Conclusions 

The application of SPF for truck body panels is 

well under way. An initial set of prototype parts was 

produced for “show and tell.” The ability to form 

such complex-shaped parts in lightweight aluminum 

was received very positively. However, based on the 

prototype trials and the increase in material cost, the 

cost data for the production parts were revised by 

the supplier. 

Production of additional prototype parts was 

delayed because of a lack of material. After some 

delay in obtaining the SPF material from the current 

Japanese supplier and a domestic supplier, the 

prototype forming trials were restarted. 

Trials with material supplied by the domestic 

supplier and a thinner-gauge material from the 

Japanese supplier were unsuccessful. 

During the prototype part development trials, it 

became apparent that the part could not be made 

consistently without splitting in some area because 

of the variations in the manually controlled forming 

process parameters. There was a very significant 

scrap rate. Also, some of the design radii may have 

been too sharp in view of the material thickness. 

The only way to determine if the selected part 

can be consistently made without significant 

problems is to run the part in a production process. 

A decision was made to modify the prototype tool to 

be production-capable by incorporating heating 

elements, polishing, and coating for friction 

reduction and by changing some of the design radii 

to make them better suited for the process. 

Since this decision to productionize tooling 

increases spending funds for material, the funding 

remaining for the validation testing will be reduced. 

If needed, International will fund the additional 

testing. 

Presentations 

Nirmal Tolani, Heavy Vehicle Materials 

Program Review, Oak Ridge, Tennessee, September 

14, 2005. 

113
High Strength Weight Reduction Materials FY 2006 Progress Report 

114
High Strength Weight Reduction Materials FY 2006 Annual Report 

B. Advanced Composite Support Structures 

Principal Investigator: Brian Knouff 

National Composite Center 

2000 Composite Drive 

Kettering, OH 45420 

(937) -297-9458; fax: (937) 297-9440; e-mail: bknouff@compositecenter.org 

Project Manager: Jay Batten 

Delphi Corp. 

3900 Holland Rd 

Saginaw, MI 48601-9494 

(989) 757-3895; fax: (989) 757-42950; e-mail: jay.batten@delphi.com 

ORNL Project Manager: Cliff Eberle 

(865) 574-0302; fax: (865) 574-8257; e-mail: eberlecc@ornl.gov 

Chief Scientist: James J. Eberhardt 

(202) 586-9837; fax: (202) 587-2476; e-mail: james.eberhardt@ee.doe.gov 

Field Technical Manager: Philip S. Sklad 

(865) 574-5069; fax: (865) 574-6098; e-mail: skladps@ornl.gov 

Contractor: Oak Ridge National Laboratory 

NCC/ORNL Contract No.: 4000039214 

Objective 

• Lead the rapid implementation of lightweight composite materials in Class 7/Class 8 vehicles via the 

development of advanced composite support structures, specifically chassis lateral braces, which can number 

up to six per vehicle. Mass reductions are targeted for 50%. 

Approach 

• Continue to model composite support structures using finite element analysis (FEA). 

• Include part, subsystem, and system analyses and optimizations. 

• Model failure mechanisms with progressive failure analysis (PFA). 

• Work with original equipment manufacturer (OEM) customer to organize potential supplier list for production: 

— Develop criteria with weighting factor on selecting final suppliers. 

— Assemble manufacturing team to interview potential supplier candidates. 

Accomplishments 

• Finalized design approach to composite center piece only to reduce risk at joint to primary beams. 

• Generated first model to accompany this approach. 

115
FY 2006 Annual Report High Strength Weight Reduction Materials 

• Met with OEM supplier to assemble a potential supplier list for production: 

— Assembled criteria ranking list for potential suppliers. 

Future Direction 

• Determine if OEM partners would like to continue program without DOE assistance: 

— National Composite Center (NCC) reduced the cost of the program and offered to pay for tooling costs to 

alleviate engineering budget. 

Introduction 

The purpose of this effort is to lead the rapid 

implementation of lightweight composite materials 

in Class 7/Class 8 vehicles via the development of 

advanced composite support structures. This task 

specifically addresses lateral braces; primary beams 

are being targeted for future work. The mass 

reduction target is 50% and the minimum 

requirement is 30%. The benefits of mass reduction 

in commercial vehicle applications are well known. 

They include increased fuel economy and a larger 

payload, which translate into fewer total trips and 

thus fewer vehicles on the road. This leads to less 

traffic, which aids highway safety, and decreased 

emissions. Support structures offer an opportunity 

for significant weight savings. However, this area of 

the vehicle also represents a large hurdle in terms of 

composite applications and market acceptance. 

The application that is the focus of this report is 

a Class 8 tractor lateral brace. The estimated total 

annual usage of Class 8 lateral braces could exceed 

1 million by 2007. Many are simple stamped parts 

that do not qualify for composite replacement on a 

per-component basis. However, some are heavy duty 

and can benefit greatly from a composite redesign, 

as in the present case. The application currently 

being studied has a production volume of about 

50,000 per year. The total weight of the current 

design is 24 kg. A minimum weight savings of 30% 

is 7.2 kg per part. There are normally two of these 

braces per vehicle, for a minimum potential weight 

savings of 14.4 kg (31.7 lb) per vehicle. 

To facilitate this work and future endeavors, 

advanced FEA and PFA softwares are being 

employed and developed. The FEA software 

contains a unique optimization algorithm that allows 

for the rapid optimization of such design variables as 

part thickness, fiber angle, fiber type, and even part 

shape while minimizing mass, strain, or even cost. 

PFA, although still in development for 

composites, provides the ability to model composite 

behavior in durability and fatigue situations. 

Modeling 

The design scenarios previously chosen were 

narrowed down to one: a composite center piece 

attached to current end brackets. Although this 

scenario does not offer the biggest mass reduction, it 

was chosen to reduce the risk at the joint to the 

primary beams. 

The first approach considers the shape of the 

center piece. Figure 1 shows the designable zones 

that were inserted in the composite center piece to 

reduce the thickness from the end bracket joint 

(where it is needed) to the center section (where it is 

not needed). 

Figure 1. Subsystem FEA model showing thickness reduction zones between green and blue areas in composite 

center piece. 

116
High Strength Weight Reduction Materials FY 2006 Annual Report 

Figure 2 shows the results. The end section 

required a composite thickness of about 26 mm, 

while the center section narrowed down to 1 mm. 

Figure 2. Results in first approach showing the thickness 

reduction. 

The resulting normal stresses shown in Figure 3 

are acceptable with maximums of +119 MPa and 

−139 MPa. However, the resulting shear stresses 

may be of some concern with maximums of +48 

MPa and −59MPa, as shown in Figure 4. 

Figure 3. Resulting normal stresses show maximums of 

119 MPa in tension and −139 MPa in compression, which 

are well below the predicted failure stresses. 

Production Readiness 

To prepare for production, a meeting with the 

OEM supplier was conducted to initialize a potential 

supplier list for production of this product. Table 1 

Figure 4. The resulting shear stresses of +48 MPa and 

−59 MPa are actually of more concern. 

Table 1. Criteria and weighting factor for ranking suppliers 

1. Process experience (thermoset only) 

• RTM/RIM/VARTM w/fabric or performs 

• Filament winding/compression molding 

• Prepreg 10 

— Unilayers (laminate) 

— Woven fabrics 

• Hybrid insert molding 

2. Cost transparencies 8 

3. Mfg. engineering support records (e.g., FMEAs, PPM 8 

record) 

4. Program management structure 7 

5. Quality control records/audits 6 

6. Commercial transportation OEM tier 1 experience 5 

7. Cost improvement plan 5 

8. Warranty policy 4 

9. Proof of customer satisfaction (awards, recognitions) 3 

10. Subcontractor list (if any) 

11. Tool support/maintenance relationships 2 

shows the criteria and weighting factor for each as 

the potential suppliers were rated. 

Conclusions 

Accomplishments for FY 2006 include the 

following: 

Modeling 

• Narrowed the design scenarios to one: 

composite center piece with current end brackets 

• Determined the thickness reduction from the 

ends to the center section 

• Finalized design plans using Genesis and 

ABAQUS for production 

117
FY 2006 Annual Report 

Production Readiness 

• Assembled manufacturing team from OEM and 

NCC personnel 

• Identified a list of nine potential suppliers 

• Developed a criteria list to rank potential 

suppliers 

Future Direction 

Modeling 

Had the program continued, the two load cases 

of racking and twisting would be used to determine 

appropriate loads applied at the same places as 

deflections to generate duplicate stress profiles. 

Those loads would then be used to optimize 

composite design in Genesis (we found previously 

that enforced displacements did not work for 

optimization). Those results would then be 

transferred to ABAQUS to optimize the metal corner 

flange design for contact stresses. March 7 was the 

expected deadline to have the finalized design 

delivered to potential suppliers. 

Production Readiness 

A program to visit the main suppliers was being 

put together to give them the overall design and 

answer questions they may have had. The list would 

be narrowed to four suppliers who would give 

request for quotation presentations to the OEM 

customer. 

High Strength Weight Reduction Materials 

Each supplier would be given 90 minutes to 

present while focusing on addressing the criteria. 

The presentations were to be scheduled April 1–11. 

On April 13, the manufacturing team would meet 

with the OEM to reduce the candidates to four. The 

design team would then visit each of the four 

candidates to answer specific detailed questions for 

production. The manufacturing team would then 

visit each supplier for its request for quotation final 

presentations the week of June 12. 

Proposal for Continuation 

NCC developed a plan that was presented to the 

OEM for continuing this program. To lessen costs, 

the PFA part of the design was eliminated and the 

focus was on safety factors in the elastic design. 

Previously, the main OEM had committed to 

funding up to $200,000 in tooling costs. To help 

alleviate engineering expenses for the OEM, NCC 

offered to pay for up to $200,000 for tooling costs 

using Ohio state capital funding. The requirement, 

however, was that all tooling and prototyping be 

performed in the state of Ohio. Unfortunately, the 

OEM was not able to transfer tooling budget to 

engineering, and the program has now been 

indefinitely suspended. 

118
High-Strength Weight Reduction Materials FY 2006 Progress Report 

C. New-Generation Frame for Pickup/Sport Utility Vehicle Application 

Principal Investigator: Curt A. Lavender 

Pacific Northwest National Laboratory 

P.O. Box 999, M/S K2-03, Richland, WA 99352 

(509) 372-6770; fax: (509) 375-4448; e-mail: curt.lavender@pnl.gov 

Kurt M. Knop 

DaimlerChrysler Corporation 

14250 Plymouth Road 

Detroit, MI 48227 

(313) 659-6076; fax: (313) 635-5221; e-mail: kmk4@daimlerchrysler.com 

Chief Scientist: James J. Eberhardt 

(202) 586-9837; fax: (202) 587-2476; e-mail: james.eberhardt@ee.doe.gov 

Field Technical Manager: Philip S. Sklad 

(865) 574-5069; fax: (865) 574-6098; e-mail: skladps@ornl.gov 

Contractor: Pacific Northwest National Laboratory 

Contract No. DE-AC06-76RLO 1830 

Objective 

• Evaluate the design of an optimized hybrid materials frame that represents a new generation of pickup/sport 

utility vehicle (PU/SUV) frame applications and vehicle architecture. 

Approach 

• Apply high-risk manufacturing and design methods to the PU/SUV frame to reduce mass while meeting cost 

goals consistent with a high-production vehicle. 

Accomplishments 

• Established performance, packaging, and weight targets for the second iteration of the new-generation frame, 

the “next-generation frame” (NGF). 

• Created a design for the NGF that projects greater weight reduction and a decrease in the number of parts 

compared with the current steel baseline frame. 

• Created a computer-aided engineering (CAE) model of the NGF and evaluated impact; noise, vibration, and 

harshness (NVH); and durability of the NGF. 

• Completed CAE and design iterations to meet requirements for a DCX 5-Star crashworthiness rating. 

• Created a complete bill of materials for the prototype and initiated procurement. 

• Assembled the prototype frame into a full-size vehicle and road tested it at the DCX Auburn Hills Technical 

Center. 

• Performed extensive analysis of the crush tip mechanical properties (the results will be the subject of a 

presentation co-authored by DCX at the Society of Automotive Engineers annual meeting). 

119
FY 2006 Progress Report High-Strength Weight Reduction Materials 

Future Direction 

• Have the frame evaluated by DCX for torsion 

• In support of DCX crush tip testing, evaluate the effect of hydroforming on tip mechanical properties. 

Introduction 

Increased consumer demand for PU/SUVs has 

resulted in increased fleet fuel consumption. The 

fuel demand for this class of vehicle has exceeded 

that of passenger automobiles and now consumes 

approximately 27% of the oil used in the United 

States.1

 The objective of this project is to explore 

manufacturing methods and materials to reduce the 

mass of the SUV/PU frame, thereby reducing fuel 

consumption for this class of vehicle. 

During the second quarter of FY 2003, DaimlerChrysler completed vehicle testing at the DCX 

Proving Grounds using an SUV/PU platform 

equipped with a hybrid frame. Results of the accelerated testing have proved that (1) the hybrid frame 

design had sufficient strength and durability to meet 

the vehicle performance requirements; and (2) the 

frame was probably somewhat overbuilt and heavier 

than required, even with a substantial weight savings 

from the current baseline steel frame. 

The next phase of the project will evaluate the 

use of a lighter frame, the NGF. The NGF uses a 

CAE approach and higher-risk manufacturing technologies. The NGF is projected to weigh less than 

the previously tested new-generation frame and require 35% fewer components. 

Approach 

A CAE model of the NGF will be created and 

design iterations performed to meet the NVH, impact, and durability requirements for a DCX 5-Star 

rating. A prototype of the frame will be fabricated 

and evaluated by frame flexure and road tests. 

Progress 

CAE analyses of the frame are complete and satisfy all of DCX requirements for 5-Star crashworthiness, NVH, and durability. The frame has been 

completed and delivered to DCX and Magna for 

full-scale automotive testing. Figure 1 shows the 

frame being fabricated in the assembly fixture. 

Figure 1. NGF components and assembly fixture before 

final assembly. 

The completed frame was assembled into a 

Dodge Durango and evaluated using a vehicle simulator at the DCX Auburn Hills Technical Center. 

The frame successfully completed a satisfactory lifetime. Figure 2 is a photograph of the Dodge Durango undergoing vehicle testing with the NGF. 

Figure 2. Dodge Durango undergoing vehicle testing 

with the NGF. 

Future Direction 

All vehicle testing of the frame has been completed and is being evaluated by DCX and PNNL 

staff. 

120
High-Strength Weight Reduction Materials FY 2006 Progress Report 

Reference 

1. Light Duty Vehicle Trends: 1975 through 2004, 

EPA 420-R-04-00, United States Environmental 

Protection Agency, Washington, D.C.,April 

2005. 

121
FY 2006 Progress Report High-Strength Weight Reduction Materials 

122
High Strength Weight Reduction Materials FY 2006 Annual Report 

D. Advanced Superplastic Forming Development for Heavy Vehicle Structures 

Principal Investigator: Curt A. Lavender 

Pacific Northwest National Laboratory 

P.O. Box 999, M/S K2-03, Richland, WA 99352 

(509) 372-6770; fax: (509) 375-4448; e-mail: curt.lavender@pnl.gov 

Don Trettin 

PACCAR Technical Center 

Farm to Market Road 

Mount Vernon, WA 

(360) 757-5241; fax: (360) 757-5370; e-mail: don.trettin@paccar.com 

Nirmal Tolani 

2911 Meyer Road 

PO Box 1109 

Fort Wayne, IN 46801-1109 

(260) 461-1238; fax: (260) 428-366; e-mail: nirmal.tolani@nav-international.com 

Chief Scientist: James J. Eberhardt 

(202) 586-9837; fax: (202) 587-2476; e-mail: james.eberhardt@ee.doe.gov 

Field Technical Manager: Philip S. Sklad 

(865) 574-5069; fax: (865) 574-6098; e-mail: skladps@ornl.gov 

Contractor: Pacific Northwest National Laboratory 

Contract No. DE-AC06-76RLO 1830 

Objective 

• Evaluate the use of superplastic forming (SPF) for heavy vehicles 

Approach 

• Focus on demonstrating the technology, using mutually agreed-on truck components, with the goal of 

developing the SPF design and material property knowledge base to the point where the individual companies 

have the ability to implement it for their new vehicle designs. 

Accomplishments 

• Performed room-temperature tensile properties tests of 5083 sheet superplastically formed to a biaxial strain of 

0.45. 

• Performed superplastic tensile tests on the baseline 5083 aluminum alloy. 

• Designed a die to evaluate the elimination of secondary operations. 

• Performed initial tests on the 6013 alloy to determine if a heat-treatable alloy may be used in a low-cost SPF 

process. Results indicate that a post-SPF aging treatment will be required. 

123
FY 2006 Annual Report High Strength Weight Reduction Materials 

Future Direction 

• Form additional trays and develop an S-N fatigue curve. 

• Perform tests with the new die. 

• Develop a process for a low-cost high-strength SPF alloy. 

Introduction 

Replacement of low-strength glass-fiberreinforced plastics with aluminum in heavy vehicle 

hoods and other cab components can significantly 

reduce the weight of Class 6-8 truck components. 

Although the use of aluminum has been viewed as a 

desirable weight savings approach for some time, 

the complex shapes of aerodynamic hoods, bumpers, 

and fairings; the limited room-temperature 

formability of aluminum; and the high cost of 

forming tools have restricted its use. SPF, in the 

context of this proposal, is an elevated-temperature 

gas pressure forming technology that has been 

widely used in aerospace applications and was more 

recently introduced by General Motors (in a 

modified form) for selected aluminum automotive 

components. Advantages of SPF include 

inexpensive tooling, the ability to form complex 

aerodynamic shapes, simplified die design compared 

with traditional stamping and the opportunity for 

significant part count consolidation. Although SPF 

is traditionally viewed as a slow-forming process, 

recent advances in aluminum alloys and forming 

process procedures have reduced typical forming 

times to the point where SPF appears well suited for 

typical heavy truck production volumes. However, a 

number of technical barriers remain, including the 

ability to form Class A surfaces, the availability of 

suitable SPF sheet materials for large components, 

and the performance of SPF components and 

structures in heavy-duty truck applications. 

Approach 

The objectives of this project are to evaluate 

current production SPF capabilities and limitations 

through parallel truck original equipment 

manufacturer (OEM) demonstration components, to 

develop cost and design allowable tools that allow 

the OEMs to select proper applications for SPF, and 

to advance SPF materials and processing capabilities 

for heavy vehicle applications. The project team 

involves collaboration between two major truck 

manufacturers, International Truck and PACCAR 

Technical Center (PTC) (parent company of 

Kenworth and Peterbilt). Because of the competitive 

nature of truck manufacturing, the project will focus 

on demonstrating the technology using mutually 

agreed-on truck components; however, it will have 

the goal of developing the SPF design and material 

property knowledge base to the point where the 

individual companies have the ability to design and 

implement it for their new vehicle designs. 

Progress 

International and PTC have identified candidate 

components for SPF prototypes. International has 

initiated manufacturing of the components, and PTC 

is completing a stress analysis of the selected 

component. 

To facilitate design of the components, Pacific 

Northwest National Laboratory (PNNL) 

superplastically formed samples of the 5083 

aluminum alloy sheet to be used, and measured 

tensile properties normal and parallel to the rolling 

direction after a biaxial strain of 0.45 at 850°F and 

1.3×10−3

 s−1 strain rate. Room-temperature tensile 

samples were removed from the long dimension of 

the formed tray (shown in Figure 1) and tested 

according to ATSM E8. Additional samples were 

subjected to the 850°F thermal cycle only and 

compared with the tensile tests from formed trays. 

The results of the tensile tests conducted in the 

rolling direction indicate that after forming to 0.45 at 

1×10−3

 s−1 at 850°F ,the sheet properties were similar 

to the as-annealed properties. Yield strength, 

ultimate strength, and elongation were 23,000 psi, 

44,000 psi and 32% for annealed materials and 

22,000 psi, 43,000 psi, and 31% for formed 

materials. 

The results of tensile tests conducted transverse 

to the rolling direction indicate that after forming to 

0.45 at 1×10−3

 s−1 at 850°F, the sheet properties were 

124
High Strength Weight Reduction Materials 

Figure 1. Sample SPF part used to characterize 

tensile properties from 0.047-in. 5083 aluminum

sheet. 

similar to the as-annealed properties. Yield strength, 

ultimate strength, and elongation were 23,000 psi, 

42,000 psi, and 30% for annealed materials and 

21,000 psi, 41,000 psi, and 29% for formed 

materials. The results of the tensile tests are 

summarized in Table 1. 

The 5083 alloy used for SPF generally has a 

very fine grain structure, developed by high levels of 

cold work, that results in a low flow stress under 

controlled temperature and strain rate conditions. 

The material supplied to PNNL was tested in tension 

at room temperature and was found to have a 57,000 

psi yield and 66,000 psi ultimate strength and an 

elongation of 10%. The high-strength sheet material 

can be beneficial prior to SPF because of high 

scratch and dent resistance forming; however, it can 

result in higher material costs if the sheet cannot be 

coiled, particularly in thick gage. 

The sheet material was tested for superplasticity

at 850 and 950°F and exhibited tensile ductility and 

strain-rate sensitivity, m-value, expected from fine 

grain superplasticity in the 5083 alloy.

1

 The samples 

were tested using constant strain rate as 

approximated by an exponentially increasing 

crosshead velocity 2 at 5×10−4

, 1×10−3, and 5×10−3 

s

−1

. At a strain of 0.15, the crosshead velocity was 

increased by 20% and the stress increase was used to 

determine the strain-rate sensitivity, m-value, using 

the relationship 

m = log(σ 2 /σ 1 ) / log(ε&2 / ε&

1 ) . 

A typical stress-strain curve derived for an SPF 

tensile test has been included here for illustration 

4500 

4000 

3500 

3000 

2500 

2000 

True Stress, psi 1500 

1000 

500 

0 

0 0.2 0.4 0.6 0.8 1 1.2 1.4 1.6 

True Strain 

0.001/s Strain Rate 0.0005/s Strain Rate 0.005/s Strain Rate

FY 2006 Annual Report

purposes as Figure 2. The superplastic tensile test 

results are summarized in Table 2. 

Figure 2. Typical superplastic tensile test curve 

performed at constant strain rate with a 20% velocity 

bump at 0.15 strain to determine strain rate sensitivity, 

m. The results shown in Figure 2 were from tensile tests 

performed at 850°F from 0.047-in. 5083 aluminum

sheet. 

The m-values ranged from a low of 0.33 at 

850°F at 5×10−3

 s−1 to a high of 0.55 at 950°F at 

5×10−4

 s−1. The elongation was relatively insensitive 

to testing conditions and was 230 to 300% except 

for the 850°F 5×10−3

 s−1 test condition, where 

elongation values dropped to below 200%. 

Consistent with the high strain rate sensitivity of 

superplastic materials, the initial flow stress of the 

sheets was dependent upon test conditions and 

ranged from a low of 500 psi at 5×10−4

 s−1 and 

950°F to a high of 3500 psi at 5×10−3 s−1 and 850°F. 

Recent developments in the automation of SPF have 

significantly reduced the cycle time required to load, 

form, and unload an SPF part. The newly reduced 

cycle time and transfer presses may enable the use of 

higher-strength alloy that has a heat treatment 

response analogous to “paint baking” in steels. 

Therefore, a brief study was initiated to determine if 

alloy 6103, which has higher strength and lower 

susceptibility to stress corrosion cracking, could be 

used. The study used hardness data to determine if a 

solution heat-treated sheet subjected to a rapid 

heatup to 850 and 950°F could exhibit an aging 

response, and if a sheet formed at 850 to 950°F 

125 

FY 2006 Annual Report High Strength Weight Reduction Materials 

Table 1. Summary of room-temperature tensile tests performed on samples machined from 

superplastically formed trays made from 0.047-in. 5083 aluminum sheet; all trays were 

formed at an average strain rate of 1.3×10−3 s−-1 and 850°F with a total cycle time (heat up to 

cool down) of 12 minutes 

Yield 

strength Ultimate 

(0.2% offset) strength Elongation 

Sheet condition Orientation (1000 psi) (1000 psi) (%) 

As-received sheet Rolling direction 57 65 10 

Transverse 57 67 9 

Annealed at Rolling direction 23 44 32 

850°F for 12 

minutes Transverse 23 42 30 

Formed tray Rolling direction 23 42 30 

Transverse 21 41 29 

Table 2. Superplastic tensile properties measured for the 0.047 inch 5083 aluminum sheet. 

Strain Initial flow 

Temperature 

(°F) 

rate 

(s−1) Orientation m-value 

stress 

(psi) 

Elongation 

(%) 

850 5×0−4 Rolling 0.54 1200 300 

Transverse 0.50 1200 300 

1×10−3 Rolling 0.49 1800 240 

Transverse 0.45 1800 240 

5×10−3 Rolling 0.34 3300 190 

Transverse 0.33 3500 180 

950 5×10−4 Rolling 0.55 500 260 

Transverse 0.51 540 250 

1×10−3 Rolling 0.50 1000 250 

Transverse 0.43 1000 230 

5×10−3 Rolling 0.35 2000 260 

Transverse 0.37 2100 230 

and removed rapidly from a die could have a post- hardness without a water quench. Figure 3 shows 

SPF age response without a water quench. the aging curves. 

Initial results showed that the hardness of a The favorable results found in the hardness 

sheet heated and held at 850 and 950°F, even for testing were followed up with tensile tests and, as 

only 1 minute, was less than 10 HRB, indicative of predicted, tensile yield strengths of slow-cooled 

a fully annealed sheet. This suggests that it is and aged samples were in excess of 34,000 psi. 

unlikely that the rapid heatup and short cycle time This significant finding will allow for additional 

associated with an automated SPF process could weight savings. A stress analysis for the PTC 

result in a “paint bake” for a solution heat-treated demonstration part showed that if an alloy with 

6013 sheet. However, 6013 sheets heated to 850 34,000 psi can be made cost-effectively, the mass 

and 950°F, removed quickly from a furnace and savings over external sheet moulding compound 

air- cooled, exhibited an age response when held panels can be in excess of 40%, compared with 

at 350°F. The hardness achieved in the air-cooled approximately 20% for commercial 5083. 

sample was less than that of a sample that was A survey of the existing literature regarding 

water-quenched; however, it did indicate that a 6××× was performed. The data 3,4

 indicate that the 

post-SPF aging treatment could result in a higher 6013/6111 alloy could be made superplastic by 

126
20 

30 

40 

50 

60 

1.2 

High Strength Weight Reduction Materials FY 2006 Annual Report 

80 

70 3000 

Hardness Rockwell B Scale 

2500 

10 

0 

0 200 400 600 800 

Time, minutes 

1000 1200 1400 1600 

950F Water Quench 950F Air Cool 850F Air Cool 

Figure 3. The age hardening response, measured in 

Rockwell B scale, of the aluminum alloy 6013 at 350°F 

after being subjected to air-cooling and water 

quenching at temperatures typical of SPF. 

controlled thermomechanical processing. 

However, the processes used required extended 

aging times that would be prohibitively costly for 

heavy vehicle sheet products. A variant was 

developed at PNNL that simply uses cold work of 

annealed or heat-treated sheet to produce the fine 

grain size required from SPF. The grain sizes 

achieved at PNNL were less than ASTM size 11 

(Figure 4) and exhibited superplastic elongations 

of greater than 200% at flow stresses typical of 

commercial fine-grained 5083. The 6013 samples 

had m-values greater than 0.4, indicating a 

relatively diffuse necking material that should 

produce high-quality parts. See Figure 5 for 

superplastic tensile test results for a fine-grained 

6013. 

A thermomechanical process developed at PNNL 

was used to produce sheet materials large enough 

to form an SPF component and allow for the 

evaluation of tensile properties using standard E8 

tensile specimens. For the tensile strength 

500 

0 

0 0.2 0.4 0.6 0.8 1

True Strain 

Figure 5. Superplastic tensile test performed on 6013 

cold-rolled 60% (solid line) and 90% (dashed) line 

showing a relative insensitivity to processing conditions. 

Tested at 5H10−3

 s−1 at 960°F. 

evaluation, PNNL used the tray die shown in 

Figure 1 and removed samples as shown in 

Figure 6. 

To simulate production cycles required for 

heavy production, the trays were formed using a 

total cycle time limit of 15 minutes (including 

heating and cooling) and a strain of 0.7 (100% 

elongation). The typical cross-section of a formed 

tray is given in Figure 7. In order to achieve the 

maximum strength levels found in the previous 

heat treatment study, the trays were formed above 

the solid solution temperature of 950°F. 

Typically, in heat treatable aluminum alloys, 

maximum strengths are achieved by water 

quenching from a sufficiently high temperature to 

produce a metastable solid solution that forms fine 

(submicron) precipitates during a low-temperature 

thermal treatment (aging). For large sheet parts, 

such as those required for heavy vehicles, water 

quenching is not possible as a result of distortion 

and costly straightening and re-strike operations. 

True Stress, psi 2000 

1500 

1000 

Figure 4. Typical ASTM 11 grain size found in 6013 

alloy subjected to 60% or greater reduction of area at 

room temperature. 

Figure 6. Tensile sample layout for 6××× alloy SPF 

trays formed at PNNL. 

127
FY 2006 Annual Report High Strength Weight Reduction Materials 

Figure 7. Typical formed 6××× alloy tray representing 

100% elongation (0.7 strain). 

Therefore, the trays were formed at 1022o

F and 

subjected to three cooling rates—water quench, 

forced air, and static air—and then aged at 350°F 

for 8 hours. Tensile results are summarized in 

Table 3. 

As expected, the tensile properties measured 

from the trays were found to be a strong function 

of the post-SPF cooling rate. The maximum 0.2% 

offset yield strength of 47,000 psi was found in the 

tray subjected to the water quenching, and 

strengths of 44,000 and 42,000 psi were found in 

the forced-air and static air-cooled trays, 

respectively. All strength levels measured were in 

excess of the target value established by PACCAR 

and indicate that a low-cost high-strength SPF 

6××× alloy is viable for heavy vehicle production. 

Table 3. Tensile values from superplastically formed 

trays subjected to different cooling rates from the SPF 

temperature of 1022°F 

Quench condition Yield UTS Elongation 

from forming 

temperature 

strength 

(1000 psi) 

(1000 

psi) 

(%) 

Water 44 59 18 

Forced air 42 56 15 

Air cooled 40 53 12 

Future Direction 

The development of data for use in SPF 

component design will continue. Fatigue data from 

superplastically formed 5083 aluminum sheets 

will be determined. 

The evaluation of a low-cost, high-strength 

alloy process for high-rate SPF will continue with 

the development of a low-cost thermomechanical 

process. 

References 

1. J. S. Vetrano et al., “Superplastic Behavior in 

a Commercial 5083 Aluminum Alloy,” 

Scripta Metallurgica 30, 565–570 (1994). 

2. K. I. Johnson et al., “The Effect of Specimen 

Geometry on the Accuracy of Constitutive 

Relations in Superplastic 5083 Aluminum 

Alloy,” in Proceedings of the 1994 

International Conference on Superplasticity in 

Advanced Materials ((ICSAM-94), Moscow, 

Russia, May 24–26, 1994. 

3. L. P. Troeger and E. A. Starke, “Particle 

Stimulated Nucleation of Recrystallization for 

Grain Size Control and Superplasticity in an 

Al-Mg-Si-Cu Alloy,” Mat. Sci. and Engr. 

A293, 19–29 (2000). 

4. R. Kaibyshev et al., “Effect of Liquid Phase 

on Superplastic Behavior of a Modified 6061 

Aluminum Alloy,” Scritpa Metallurgica 47, 

569–575(2002) . 

128
______________________________________________________________________________ 

High Strength Weight Reduction Materials FY 2006 Progress Report 

E. Development of Magnesium for Heavy Vehicle Powertrain Components 

Principal Investigator: Darrell R. Herling 

Pacific Northwest National Laboratory 

P.O. Box 999, Richland, WA 99352 

(509) 376-3892; fax: (509) 376-6034; e-mail: darrell.herling@pnl.gov 

Chief Scientist: James J. Eberhardt 

(202) 586-9837; fax: (202) 587-2476; e-mail: james.eberhardt@ee.doe.gov 

Field Technical Manager: Philip S. Sklad 

(865) 574-5069; fax: (865) 574-6098; e-mail: skladps@ornl.gov 

Contractor: Pacific Northwest National Laboratory 

Contract No.: DE-AC06-76RL01830 

Objective 

• Develop magnesium (Mg) composite casting technology and associated low-cost tooling for manufacturing of 

ultra-lightweight heavy vehicle powertrain components. The project is focused on three development tasks: 

— Assess the potential to use Mg for vocational and long-haul trucks through technology development, 

evaluation of full-scale prototype components, and economic analysis. 

— Develop advanced Mg materials and processing capabilities for heavy vehicle applications. 

— Develop a low-cost shape-forming process that combines casting and in situ compositing for metal matrix 

composite (MMC) components. 

Approach 

The following is a project outline and general approach. The underlying focus is on process development and 

economical manufacturing methods for Mg and its composites. The approach and cost analysis will consider 

product design, material-specific characteristics, and component performance. 

• Address technical challenges for the use of Mg: 

— Develop high-strength, low-cost Mg MMC—PNNL/supplier. 

— Current Mg alloys will have limited application for heavy vehicles. 

— High-strength Mg MMC can improve strength while minimizing cost. 

— Increase durability; fatigue life is more demanding than in automobiles and operating temperature/creep 

resistance are needed. 

— Develop economical manufacturing/casting/compositing method. 

 Develop stir-cast Mg process. 

 Demonstrate hybrid lost-foam/pressure infiltration casting. 

• Apply Mg to heavy vehicles: 

— Determine how and where it is best to use Mg—PNNL/original equipment manufacturer (OEM). 

 Develop design guidelines and properties database. 

— Work to reduce manufactured part cost—OEM/supplier/PNNL. 

 Address issues with integration of Mg into vehicle manufacturing process. 

129 
FY 2006 Progress Report High Strength Weight Reduction Materials 

Accomplishments 

• Designed, fabricated, and installed a new rapid compositing mixer at PNNL specifically for Mg MMC 

composites. 

• Initiated experimental trials for stir-cast Mg-MMC production and modeled lost-foam casting molds. 

Future Direction 

• Develop the compositing ability to in situ cast and mix an Mg composite casting. 

• Evaluate a modified lost-foam casting process that combines casting and compositing into one process step. 

Introduction 

Ultra-lightweight, high-performance Mg MMCs 

can improve fuel utilization while increasing 

component durability for heavy vehicles. As a rule 

of thumb, lightweighting can contribute to an 

increase of roughly 6–7% in fuel economy for every 

10% reduction in mass. An estimated 3000 lb of 

mass can be targeted for reduction in the heavy 

vehicle powertrain and chassis. There is the potential 

for roughly 200–500 lb of weight savings through 

redesigning and replacing specific aluminum and 

cast iron components with Mg MMCs. 

Downsizing of ancillary fixtures can result in 

additional mass and fuel savings. Increased 

component durability in applications, where wear is 

a factor, can also lead to improved vehicle operating 

costs. This is an incentive for fleet owners and 

independent truckers to embrace the new 

technology. 

The objective of this project is to enhance the 

properties of Mg alloys by creating a Mg-ceramic 

composite. This class of material is an MMC in 

which ceramic particles or short fibers are added to 

the metal matrix. The resulting composite has very 

high specific strength and stiffness and significantly 

improved wear resistance. Thin cast sections are still 

possible; and, in the case of Mg, noise, vibration, 

and harshness (NVH) and load carrying capacity 

improve beyond the basic alloy. This is especially 

beneficial for vocational trucks, where load capacity 

and noise/vibration are two of the more common 

design problems facing engineers. 

Approach 

Development of low-cost Mg alloys and 

processing technologies is a major technical hurdle 

to widespread application of Mg in the heavy 

vehicle manufacturing industry. Therefore, the 

approach taken for this project includes active 

participation by a truck OEM, as well as by 

suppliers where appropriate. The philosophy of this 

teaming approach is to ensure meaningful 

technology transfer of enabling and new technical 

developments. 

This project is a collaborative effort with Mack 

Trucks, Driveline Division. The project goal is to 

develop Mg composite casting technology and 

associated low-cost tooling for manufacturing ultralightweight heavy-vehicle powertrain components. 

The project scope is broken down into three tasks: 

1. Assess the potential to use Mg for vocational and 

long-haul trucks through technology 

development, evaluation of full-scale prototype 

components, and economic analysis 

2. Develop advanced Mg materials and processing 

capabilities for heavy vehicle applications 

3. Develop a low-cost shape-forming process that 

combines casting and in situ compositing for 

MMC components. 

Metal Compositing Method 

Stir-casting techniques are currently the most 

common commercial method of producing 

aluminum MMC materials. This approach uses 

mechanical mixing of the reinforcement particulate 

into a molten metal bath. The resulting metal 

composite is cast, typically, by low-pressure die or 

sand casting. This process is effective in making 

high-quality parts. The process, coupled with a lowcost MMC feedstock material, can be an economical 

manufacturing method. However, combining the 

compositing and casting processes into a single 

process reduces the cost of producing an MMC 

component by an estimated 6%. 

We have initiated an experimental investigation 

of the stir-cast Mg MMC production method as one 

130 
High Strength Weight Reduction Materials FY 2006 Progress Report 

of the two technical development thrusts of this 

project. 

A hybrid system that consists of combining 

evaporative pattern (lost-foam) casting techniques 

and low-pressure infiltration to create the metalceramic composite during the shape forming 

(casting) operation is currently being modeled. This 

second process, though a higher technical risk for 

success, would significantly reduce componentmanufacturing cost. We expect it to allow easy, 

selective reinforcement of cast components because 

the reinforcement particles would be placed only 

where they would be of the greatest benefit to part 

performance. The key to achieving this approach is 

the ability to co-mix the second-phase particulate 

with the foam beads prior to expanding into the 

mold pattern, followed by a controlled burnout rate 

of the pattern during casting and infiltration of that 

particulate by the liquid metal front. 

Once this process is modeled and the casting 

parameters are estimated, various part geometries 

will be cast in order to validate the process. Initially, 

tensile bars will be cast for mechanical property 

evaluation. 

Magnesium MMC Development 

In order to develop proper Mg MMC materials 

for powertrain components, it is necessary to 

investigate which alloy-reinforcement combinations 

will work best to achieve the mechanical 

performance goals as well as the environmental 

performance goals. To accomplish this task, a stircast metal composite mixer, shown in Figure 1, has 

been designed by PNNL and MC-21 Inc., 

constructed, and installed at PNNL. 

The micro-mixer’s design is based on MC-21’s 

rapid mixing technology for aluminum MMCs, with 

provisions to safely melt Mg. This technology is 

unique in that it is a highly efficient means of 

mixing together (compositing) ceramic and molten 

metal, resulting in a metal composite material. The 

unit is capable of melting, mixing, and casting 10 kg 

of Mg or aluminum metal in a single batch. With 

this capability, multiple combinations of alloyreinforcement chemistry can be produced and 

evaluated. 

Figure 1. New PNNL rapid mixing and casting 

system. 

Conclusions 

Magnesium MMCs have outstanding specific 

strength and stiffness properties as well as NVH 

dampening characteristics—greater than Mg alloys 

alone. They also have enhanced wear and creep 

resistance over unreinforced alloys. One drawback 

to Mg and its MMC is that the supply base has all 

but disappeared in North America, requiring OEMs 

and specialty/prototype casters to assume production 

roles. The focus of this project is to develop 

manufacturing technologies that allow for the 

production of low-cost Mg MMC components. The 

intent is to allow an OEM to apply the low-cost 

process or transfer the technology easily to a 

supplier. The development of Mg composite casting 

technology will include low-cost tooling approaches 

for manufacturing of ultra-lightweight heavy-vehicle 

powertrain components, as well as the development 

of a low-cost shape-forming process that combines 

casting and in-situ compositing for MMC 

components. 

131 
FY 2006 Progress Report High Strength Weight Reduction Materials 

132 
High Strength Weight Reduction Materials FY 2006 Progress Report 

F. Lightweight Stainless Steel Bus Frame―Phase III 

Principal Investigator: J. Bruce Emmons 

Autokinetics Inc. 

1711 West Hamlin Road, Rochester Hills, MI 48309 

(248) 852-4450; fax: (248) 852-7182; e-mail: jbemmons@autokinetics.com 

Chief Scientist: James J. Eberhardt 

(202) 586-9837; fax: (202) 587-2476; e-mail: james.eberhardt@ee.doe.gov 

Field Technical Manager: Philip S. Sklad 

(865) 574-5069; fax:(865) 574-6098; e-mail:skladps@ornl.gov 

Contractor: Autokinetics Inc. 

Contract No.: 4000010114 

Objectives 

• Investigate and demonstrate the mass-saving potential of ultra-high-strength stainless steel as applied to the 

structure and chassis of a full-size urban transit bus. 

• Finalize design and analysis and build a full-scale prototype of the body structure and chassis. 

• Investigate all of the fundamental feasibility issues related to the structure and chassis: 

1. Fabricate and test large lightweight stainless steel sandwich panels 

2. Fabricate roll-formed, high-strength stainless steel sections 

3. Test the feasibility of lightweight stainless steel cantilever seats 

4. Design and fabricate lightweight stainless steel independent suspension 

5. Integrate the traction motors into the suspension design 

Approach 

• Execute the basic body structure, including the floor/roof sandwich panels, pillar assemblies, longitudinal rails, 

and suspension subframes. 

• Choose prototyping techniques that emulate the intended production process as closely as possible to aid in developing robust but cost-effective manufacturing techniques essential to meeting the objectives of the project. 

• As computer-based design and analysis details of the bus develop, conduct hands-on physical experimentation 

in parallel to support the concepts and methods. 

Accomplishments 

• Installed exterior side panels. 

• Made progress on steering system components. 

• Received all window glass. 

• Installed windows. 

• Completed emergency exit windows. 

• Completed prototype driver’s station (mock-up). 

• Received batteries. 

133 
FY 2006 Progress Report High Strength Weight Reduction Materials 

• Installed batteries. 

Future Direction 

• Complete the fabrication and installation of suspension, steering, and spring components. 

• Assemble propulsion components. 

• Prototype two seats. 

• Test structure. 

Introduction 

Advanced-technology transit bus concepts have 

made significant advancements in terms of light 

weight and fuel economy. However, these gains 

have come at the expense of higher manufacturing 

costs. In spite of attempts to use life-cycle costs to 

justify their purchase, initial cost remains a major 

obstacle to the introduction of fuel-efficient buses. 

DOE’s OFCVT approached Autokinetics in an 

attempt to solve this problem and asked Autokinetics 

to develop concepts for a lightweight urban transit 

bus based on the use of high-strength stainless steel. 

In the passenger car field, Autokinetics had developed structural and manufacturing techniques for the 

cost-effective use of stainless steel in spaceframes 

and suspensions. OFCVT wanted to determine 

whether this approach could be applied to transit 

buses as well. 

The program was structured in three phases: 

• Phase I—Initial Concept Development 

• Phase II—Concept Verification and Initial 

Design 

• Phase III —Final Design and Prototyping of 

Body and Chassis 

Phases I and II have been successfully completed. Phase III will result in a full-size body structure and suspension that will be tested statically and 

dynamically. The development of an optimized hybrid powertrain and other vehicle systems will be 

addressed in a separate project. 

This project was unusual in that no formal mass 

or cost targets were given. The object was to save as 

much mass and cost as possible. 

Current State of Progress 

With the installation of exterior side panels and 

glass, the overall body structure is now nearly complete. Much of the remaining work will focus on the 

final design and fabrication of suspension components. Relatively few tasks remain, and it is expected 

the project will be completed by mid-2007. 

Exterior Side Panels 

The body side panels (Figure 1) were assembled 

to the basic body structure during this reporting period. Because the side panels are intended to be an 

integral part of the overall structure, it was necessary 

to install them with their material in tension. This 

requirement was complicated by the fact that the 

0.030-in. stainless steel sheet material to be used 

displayed pronounced waves due to coil-set. A special technique was developed to position and pretension the side panels as they were spot-welded to 

the structure. This technique worked very well, was 

straightforward to accomplish, and resulted in very 

little surface deviation in the installed side panels. 

Steering System Components 

Additional progress was made in the steering 

system. Various communications and meetings took 

place with Pailton Steering Systems, ZF Lemforder, 

and TRW to discuss the somewhat unusual, lowfloor-compatible steering system design and necessary components. As of the end of this reporting period, all necessary components have been identified 

and are in the process of being delivered. 

134 
High Strength Weight Reduction Materials FY 2006 Progress Report 

Figure 1. Exterior side panels. Figure 2. Body side glass. 

Body Glass 

As reported previously, laminated glass was 

chosen for all glazing positions because it provides a 

number of advantages compared with tempered 

glass. During this reporting period, Fox Fire Glass 

(Pontiac, Michigan) completed fabricating and delivered the prototype glass for all side windows and 

the two-part windshield. 

The fixed-position glass was installed by Troy 

Auto Glass Inc. (Troy, Michigan), a local automotive glass installer. It is shown in Figure 2. The installation was accomplished using standard processes and materials. It is interesting to note that a 

couple of windows were installed prematurely and 

had to be removed to allow access for other operations. Troy Auto Glass removed these windows using industry-standard equipment and reinstalled 

them when appropriate. This exercise provided the 

opportunity to evaluate the glass replacement procedure. It was found that replacement of the bonded 

side glass could be accomplished quickly and efficiently using only standard equipment, materials, 

and processes. 

Emergency Exit Windows 

In addition to the two doors, five windows located throughout the bus are designated as emergency exits. One such window is shown in Figure 3. 

Since the vehicle structure incorporates bonded-inplace side glass, the five exit windows required a 

different attachment and sealing approach. These 

windows are of a frameless design and are attached 

Figure 3. Emergency exit window. 

to the structure through a standard piano- type hinge 

bonded to the upper edge of the glass. The body side 

of the hinge is bolted rather than spot-welded to aid 

replacement. A standard profile bulb seal pressed 

onto the flanges around the perimeter of the window 

opening provides the sealing function as well as protection from the raw edges of the sheet metal. 

135 
FY 2006 Progress Report 

A latch-bar, which carries the two side-position 

latches, was then bonded to the inside of the glass 

panel. This joining procedure uses the same bonding 

material as the standard glass installation. The use of 

a latch-bar allows for subassembly of the latch system, greater dimensional control, and a wide distribution of the 03566e stress imposed by the latch. 

Driver’s Station 

A prototype driver’s station mock-up was assembled to physically locate the relative positions of 

all driver station components. A plywood “buck” 

was fabricated to which the physical components 

such as the driver’s seat, steering column, and pedals 

were attached. It also allowed for the flexibility to 

reposition the various elements of the driver’s environment. This mock-up provided a basis for evaluation and was used in the development of the driver’s 

station design task covered under a separate project. 

The completed driver station derived from this 

mock-up is shown in Figure 4. 

Figure 4. Driver’s station. 

Received Batteries 

During this reporting period, Autokinetics received delivery of nine Zebra Z37 batteries from 

MES-DEA (Switzerland). Eight of these batteries 

were purchased under a DOE Systems project. One 

battery was purchased to fulfill the requirements of 

this project in place of a bank of 25 lead-acid batteries as originally proposed. The additional cost of the 

Zebra battery was absorbed by Autokinetics. It is 

felt that this more advanced battery technology will 

result in a much more meaningful first trial of the 

High Strength Weight Reduction Materials 

bus prototype. This first battery was connected to 

the vehicle control network and the motor controllers to enable testing and debugging of the prototype 

electric drive propulsion system. 

Battery Installation 

During this reporting period, a battery arrangement was developed and mounting hardware designed and fabricated to secure the batteries to the 

vehicle. Overall mass balance, length of conductors, 

and high-voltage safe practices were all important 

factors in determining the configuration and placement of the batteries and their supporting components. 

Once preliminary testing of the control network 

was completed, the battery was installed into the 

vehicle along with its charger, cooling fan, and a 

motor controller. This installation is shown in Figure 

5. 

Figure 5. Battery installation. 

Dissemination and Commercialization 

9/23/05—Meeting with ZF Lemforder. Attended 

by Paul Donnelly and Cliff Eschelman. 

11/22/05—Meeting with Solutia, Inc. Attended 

by Jay Pyper and others. 

11/30/05—Informal review meeting with Lee 

Slezak (DOE); Donald Hillebrand, Argonne National Laboratory (ANL); Tim Murphy and Jim 

Francfort, Idaho National Laboratory. 

136 
High Strength Weight Reduction Materials 

2/23/06—Meeting with Utilimaster. Attended by 

Rob Fairchild (Fairchild and Associates) and John 

Knudtson (Utilimaster) 

3/15/06—Meeting with Pailton Steering Systems. Attended by Rob Kilhefner, Andy Rose, and 

James Wiles. 

5/23/06—Meeting with City of Rochester Hills, 

MI, to discuss “Smart Zone.” Attended by Dan Casey (City of Rochester Hills). 

8/24/06—Systems project review. Attended by 

Lee Slezak (DOE) and Jules Routbort (ANL). 

Conclusions 

Autokinetics remains confident that highstrength stainless steel has the potential to achieve 

substantial mass reductions in bus structures. The 

bus body structure is now nearly complete; most of 

the identified technical risk issues have been resolved; and structural and chassis mass estimates 

remain nearly unchanged compared with early predictions. Ongoing fabrication of the physical prototype has provided concrete mass numbers, and it is 

now expected the actual mass reduction of the complete battery electric vehicle will be close to 50%. 

It is also hoped that practical commercialization 

can be achieved in the not too distant future. Low 

capital investment and an ample knowledge base are 

key enablers for this. Much has been learned thus far 

about processing and assembling of the body structure and many useful techniques have been developed. Given the relative ease of constructing this 

prototype within our own facility, it is quite apparent 

that capital requirements for commercializing this 

technology will be relatively small. 

The independent cost analysis commissioned by 

DOE supports this as well as the predicted mass and 

unit cost savings. 

FY 2006 Progress Report 

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FY 2006 Progress Report High Strength Weight Reduction Materials 

138 
High Strength Weight Reduction Materials FY 2006 Progress Report 

6. APPLICATION OF INNOVATIVE MATERIALS 

A. Application of Carbon Fiber for Large Structural Components 

Principal Investigator: Kevin L. Simmons 

Pacific Northwest National Laboratory 

P.O. Box 999, MSIN: K2-44, Richland, WA 99352 

Phone 509-375-3651; fax: 509-375-2186; e-mail: kl.simmons@pnl.gov 

Technology Development Manager: Dr. James Eberhardt 

(202) 586-9837; fax: (202) 586-1600; e-mail: james.eberhardt@ee.doe.gov 

Field Technical Manager: Philip S. Sklad 

(865) 574-5069; fax: (865) 574-6098; e-mail: skladps@ornl.gov 

Contractor: Pacific Northwest National Laboratory 

Contract No.: DE-AC06-76RL01830 

Objective 

• Develop selective reinforcement technology that can be applied to large cab components to improve specific 

stiffness and strength while reducing overall component weight. 

Approach 

• Determine how well low-cost carbon fibers can be hybridized with glass fibers to provide substantial weight 

and cost reductions in large cab components. 

• Develop a system for preforming carbon and glass fibers together that will allow components to take 

maximum advantage of the capabilities of selective reinforcement alignment and property contribution. 

• Develop models for the analysis of hybrid chopped fiber preforms and composites that allow the thermal and 

structural properties to be developed and compared with experimental analysis. 

• Perform structural testing to define the limits of applicability of the carbon/glass hybrid reinforcement 

materials to the large structures and develop guidelines for applications that may be used by original 

equipment manufacturers (OEMs). 

• Design and develop critical sub-section components of large structures to use in correlation with the predictive 

models, and to validate the structural application criteria. Determine the capability of the materials to be 

fabricated in full-scale components and determine the performance of these full-scale components in real 

application scenarios. 

Accomplishments 

• Developed short fiber composite models for predicting properties of hybrid composite structures. 

• Completed experimental testing of hybrid composite panels for model correlations. 

• Developed a method of measuring fiber angles to determine the level of anisotropy in chopped fiber 

composites. 

• Completed a set of design experiments on a new resin system for use with the hybrid composites. 

• Completed initial runs on a large test tool with short fiber composite preforms. 

139
FY 2006 Final Report High Strength Weight Reduction Materials 

• Developed a large-scale, sub-component tool specifically to mimic design attributes applicable to several 

structural and appearance components. 

• Performed initial runs on a large test tool with short fiber composite preforms at a tier-one supplier. 

• Developed hybrid fiber preforms for test tool and demonstrated initial successful molding with hybrid fiber 

and resin system. 

• Prepared samples for testing from large structural components. 

• Developed the first Class A structural carbon composite panels for truck applications. 

• Demonstrated impact performance 2–3 times better than that of standard structural composites. 

• Developed hybrid carbon composite with improved impact performance over fiberglass composites. 

• Demonstrated thin (2.5-mm) panel fabrication and resin transfer molding capability. 

• Demonstrated panel molding with varying cross-sections from 2.5 to 7 mm in a single component. 

• Performed panel molding with various filler types and loadings and selected optimum fillers for achieving a 

Class A surface. 

• Performed mechanical and impact tests on a range of panels. 

• Developed and performed process trials with a new tooling system that allowed the system to work together. 

This included improving the thermodynamic control necessary for achieving mechanical and impact 

properties. 

• Optimized the resin system using different fillers. 

• Tested and validated thickness constraints on panels for achieving structural performance on surface-coated 

thin hybrid panels. 

• Tested and validated the barrier coat and optimized it for surface profile. 

• Modeled tooling resin flows into fiber perform. 

• Evaluated the feasibility of large-scale part production. 

• Determined the cost spread among tooling approaches compatible with this system. 

• Evaluated the dimensional stability of the hybrid composite material under various environmental conditions. 

• Evaluated the paint bakeout cycle for changes in adhesion and surface appearance. 

Introduction 

The purpose of the project is to develop the design and material processing technology that will facilitate the use of low-cost hybrid glass fiber and 

carbon fiber reinforcements in large composite components, which will reduce weight and improve 

structural performance. The project will also seek to 

advance low-cost carbon fiber materials developed 

by industry by advancing the introduction of lowcost resin systems that are compatible with current 

heavy vehicle structural composites. 

Our approach was to determine the feasibility of 

hybridizing low-cost carbon fibers with glass fibers 

to provide substantial weight and cost reductions in 

large cab components. A system was developed for 

preforming carbon and glass fibers together that 

would allow the components to take maximum advantage of the capabilities of selective reinforcement 

alignment and property contribution. Models were 

developed for the analysis of hybrid chopped fiber 

preforms and composites that allow the thermal and 

structural properties to be developed and compared 

with experimental analysis. 

Structural testing was performed to define the 

limits of applicability of the carbon/glass hybrid reinforcement materials to the large structures and develop guidelines for applications that may be used 

by OEMs. We evaluated the design and development of critical sub-section components of large 

structures to use in correlation with the predictive 

140
High Strength Weight Reduction Materials 

models, and to validate the structural application criteria. The project goal was to determine the capability of the materials to be fabricated in full-scale 

components and determine the performance of these 

full-scale components in real application scenarios. 

The use of selective reinforcements with higherstiffness fibers has the potential to reduce both costs 

and weight simultaneously, even at today’s price 

points. However, their introduction is hindered by 

the lack of understanding of the fibers in existing 

processes, the cost, and the need to develop robust 

methods of preforming glass and carbon fiber materials together. In addition, the capability to provide a 

Class A surface finish is required; developing that 

capability will require the potential development of 

bonding agents and the ability to model thermal and 

structural performance of the materials in a hybrid 

system. The recent development of composite systems combining carbon fiber reinforcement with 

low-cost automotive and marine resin systems provides the opportunity for selective reinforcement of 

a broad range of structural composite components. 

Predictive Modeling 

Summary of the Modeling Approach 

A computational tool called EMTA, which uses 

the Eshelby/Mori-Tanaka approach, has been developed to predict the elastic and thermal properties 

(thermal conductivity, thermal expansion coefficients) as well as the elastic-plastic responses for 

short-fiber composites.1

 The EMTA model makes 

use of the Eshelby equivalent inclusion method in 

which fiber characteristics (inclusion shape) are accounted for through the so-called Eshelby tensor 

while the constituents’ properties and volume fractions and the interactions of fibers are taken into account through the homogenization procedure. 

EMTA can handle two types of fibers, which are 

mixed and embedded together in a resin matrix (hybrid composites). Planar orientation is assumed and 

obeys the fiber orientation distribution density of the 

form ρ(θ ) = λe −λθ

, where λ is the orientation angle 

measured with respect to the major fiber alignment 

axis (or fiber orientation distribution axis), and θ denotes a parameter governing the randomness. When 

λ is very small (θ →0) the fibers are randomly oriented, while they are aligned for large values of θ. 

Between these two limiting cases, a given value of θ

designates a fiber orientation distribution that can be 

FY 2006 Progress Report 

rather random or rather semi-oriented. EMTA can be 

used as a stand-alone code to predict the basic homogenized composite properties. It has also been introduced into ABAQUS by means of the user 

subroutine UMAT of this code for structural analyses. 

The value in developing this code is that it helps 

the designers of large structures design shapes and 

do analyses using hybrid composites. There is very 

little modeling capability that can predict properties 

for short-fiber composites. Models that combine 

both fiber types, based on randomness of each, are 

not available. 

Figure 1 illustrates an unsymmetrical hybrid 

composite made of different composite layers, which 

are not symmetrical through the thickness. These are 

of particular concern because of the mismatch of coefficients of thermal expansion (CTEs) between the 

layers, which can lead to unacceptable structural deformations. On the other hand, adjusting the CTEs 

by means of trial-and-error approaches is timeconsuming and expensive. Therefore, numerical 

methods to predict the CTEs of the composite layers 

and the governing microstructural parameters (e.g., 

fiber volume fraction, aspect ratio, orientation distribution) are helpful in the design of unsymmetrical 

hybrid composite laminates. 

Figure 1. Schematic of a hybrid laminate 

Table 1 compares the CTEs for different materials and the prediction in the composite forms. When 

the material constituent properties are combined 

with the processing orientations that a machine may 

bias, the EMTA model is able to predict the part behavior. 

Figure 2 shows the predicted deflection (displacement u3 ) resulting mainly from the mismatch 

of the in-plane CTEs (α1 ,α 2 ) between the carbon 

141
FY 2006 Final Report 

Table 1. Thermal expansion coefficients of the constituent materials and of the fiber tows filled with Hydrex 100 

α1 α 2 α 3 

(xE6/o

C) (xE−6/o

C (xE−6/°C 

) ) 

Carbona −0.7 10 10 

E-glassa 5.4 5.4 5.4 

Vinyl esterb 70 70 70 

Carbon/vinyl −0.011935 38 38 

esterc 

Glass/vinyl es- 7.7081 36.525 36.525 

terc 

a

 ASM Handbook of Composites.

2

b Based on the values of thermoset polyester given in

ref. 2. 

c EMTA prediction. 

Figure 2. Prediction of the deformed shape an dout-ofplane deflections for the analyzed hybrid panel after 

molding. 

and glass layers. The predicted deformed shape 

agrees well with the experimental shape shown in 

Figure 3. Predicted and experimental deformed configurations are presented in Figure 4 in terms of the 

spatial coordinates X, Y, and Z. In both configurations, the Z-coordinates were determined on the top 

surface of the hybrid panel. The predicted values 

correlate well with the experimental results. 

To reduce the mismatch of CTEs, the first remedy is to realize the same orientation distribution for 

the carbon and glass fiber tows. Consider, for instance, that the carbon tows’ orientations are also 

completely random. Such orientations will lead to 

the predicted variations of the in-plane CTE as a 

High Strength Weight Reduction Materials 

Figure 3. Warpage of the hybride panel after molding. 

Figure 4. Experimental and predicted deformed coordinates for the analyzed hybrid panel after molding. 

function of the fiber tow volume fraction shown in 

Figure 5. 

For a given glass tow volume fraction, the 

curves provide the corresponding carbon tow volume fraction, which must be realized so that the 

mismatch of CTEs is zero. These values were determined for the proper fiber volume fractions of the 

glass layer and of the carbon layer, as well as the fiber volume fractions of the tows. Finally, Figure 6 

shows the simulation results for the adjusted case. 

The deflection is reduced practically to zero as a result of suitable adjustments of fiber orientation and 

of the carbon tow volume fraction in the carbon 

layer. 

Preform Materials 

Experimental preform test panels have been fabricated using a variety of methods. The initial stud142
High Strength Weight Reduction Materials FY 2006 Progress Report 

Figure 5. Variations of the in-plane CTEs of the carbon 

and glass layers with the fiber tow volume fraction. 

Figure 6. Prediction of the deformed shape and out-ofplane deflection for the analyzed hybrid panel after adjustments of fiber orientation distribution and fiber volume fraction for the carbon layer. 

ies attempted to use preforms by the National Composites Center (NCC). These panels (Figure 7 show 

a typical carbon/glass hybrid panel using 3% carbon) had variations in the volume fractions of glass 

and/or carbon. The hybrid panels also had four levels of carbon fiber in addition to the glass. Carbon 

volume levels were 3%, 9%, 15%, and 45%, with 

the balance being random glass, and were made in 

two nominal thicknesses of 2.5 mm and 4.0 mm. 

There were also fiber-directed preforms of carbon 

fiber for comparing how well the properties could be 

controlled for a specific orientation. The panels were 

then formed in a resin transfer mold (RTM) 

Figure 7. Composite hybrid perform with 3% carbon fiber chop. 

with a standard polyester RTM resin from Reichhold 

Corporation. 

Fiber orientation measurements were made for 

determining the randomness factor for the EMTA 

model. Figure 7 shows a 3% carbon fiber preform 

that was used to determine the random pattern of the 

chopped material applied to the preform. 

The preform digital image was then used in a 

Scion Image software package from which fiber orientation was measured relative to the bottom edge. 

The orientation was from 0 to 180°. Figure 8 shows 

the output data from the Figure 7 panel. The data 

show very good randomness with no defining pattern. 

NCC also provided some fiber-directed preforms for evaluation. The panels were to have 0° and 

90° directed fibers. The panels used in these measurements proved to be more challenging because of 

the higher fiber volume fraction of carbon, which 

presented a darker background and little con-trast. 

The lighting was extremely important for enhancing 

the fiber tows on the surface. 

Figure 9 shows the histogram data of the oriented fiber measurements. A pattern is developed by 

the preferential orientation of the chopped fiber pattern program from the chopper gun. The histogram 

shows a bimodal distribution of fibers at approximately +45° and −60°. 

The NCC panels did not perform well for processing. The use of a low-pressure RTM through a 

low-porosity preform created a slow process time 

that was not conducive to large part applications. 

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FY 2006 Final Report High Strength Weight Reduction Materials 

Figure 8. Histogram of fiber orientation. 

Figure 9. A bimodal distribution of carbon fiber. 

This process was not deemed useful for further 

study, given the hybrid materials approach and large 

part format we were considering. The development 

of a hybrid system with a flow media core significantly increased the processing speed and created an 

improved part. Two vendors provided commercially 

available glass materials that were modified with our 

hybrid carbon fiber approach. 

The vendor-supplied materials were evaluated 

based on core thickness and core types. Several 

binder formulations for binding chopped carbon fibers to the glass and flow media material were 

evaluated, and a final formulation was chosen for all 

of the experiments. The significance of the flow 

speed was a key criterion if the end use was to be 

large structural components. The time to fill was reduced from 5–6 min down to less than 1 min for our 

model experiment. The significance of this time is 

that it allows for the tool cavity to fill with resin before the resin begins to gel and build a high enough 

viscosity that the resin will not wet out fibers and 

stops flowing in the cavity. 

Resin Evaluation 

Initial evaluations of the EMTA model used a 

traditional low-cost polyester resin system. The resin 

system did not perform satisfactorily with hybrid fiber combinations, so the program worked with 

Reichhold Chemical to come up with a low-cost hybrid resin. Current interest in the attractive properties arising from the combination of polyester and 

urethane resin chemistries has prompted investigation into efficient manufacturing methods using a 

blended polyester/urethane system. By mating this 

material to a glass/carbon hybrid fiber perform, an 

optimization of properties from all of the constituents can be achieved at a relatively low cost, especially if the laminate production can occur within a 

short cycle time. 

The resin system chosen was a hybrid resin chosen for its ability to bond to both glass and carbon 

fiber surfaces. Traditional polyester resins bond well 

to glass fibers but are substandard for bonding to 

carbon fiber. Other resin systems that bond well to 

carbon fibers exceed the price points for a low-cost 

composite system. Those systems would not be costeffective for manufacturers to implement. The hybrid system underwent a set of experiments to determine the controlling variables for processing. 

Several catalyst types were evaluated; temperature 

ranges, promoter types, and levels were investigated. 

A Minitab design of experiments (DOE) was run using a full factorial design with two levels and three 

factors. The interest in the experiments was to help 

determine which components gave the most control 

over the gel time. The factors of interest were temperature, cobalt content, and catalyst. The output 

measurements were gel time, peak exotherm time, 

and peak temperature. 

The experiments were run on a Haake Rheocord 

90 mixer using Brabender medium-shear mixing 

blades. The new hybrid resin from Reichhold was 

used. The catalyst, cobalt promoter, and temperature 

were controlled for the experiments. 

The data were collected based on time, temperature of the resin, and torque value during the mix. 

The gel time was taken as the first inflection of the 

torque curve. This is based on the torque value be144
SMC

Liquid/PreformHybrid/glass

Hybrid/Carbon

High Strength Weight Reduction Materials 

ginning to climb during resin gelling. The resin temperature also begins to increase relative to the same 

point as the gel time. The peak exotherm time is 

taken at the peak temperature that the mixture 

reaches before cooling down. 

Figure 10 is a Pareto chart that illustrates the 

standardized effects from the experiments. The 

Minitab output compares the different factors and 

combinations of effects. The Pareto chart clearly indicates that temperature has the most effect on gel 

time, followed by the cobalt content. The combined 

effect of cobalt and temperature is a key contributor 

to the gel time. The catalyst showed very little effect 

on the gel time, and the trend was for the lower catalyst content to trend higher in peak exotherm temperature. However, the difference in peak 

temperatures is not statistically valid and is not considered as a concern. 

Pareto Chart of the Standardized Effects 

(response is Gel Time, Alpha = .10) 

A: Temperat

B: Cobalt A C: Catalyst 

B 

AB 

ABC 

C 

AC 

BC 

0 10 20 

Figure 10. Pareto chart for gel time. 

This set of experiments gave us the controlling 

parameters needed for molding large structures. It 

provided the controls to slow down the reaction and 

minimize the time to cure. The experiments also 

showed what temperature expectation would occur 

during curing so that tooling can be designed to control the temperature. This would allow for part-topart processing consistency without heat buildup in 

the tooling, which would create greater variability. 

The development effort remained focused on using a carbon- and fiberglass-compatible resin system. This has proven to be a valid selection based on 

impact test data that indicate excellent performance 

with carbon fiber systems, as well as glass and other 

systems. The resin is tailored for use in an RTM 

process. The resin supplier has mentioned several 

features of the development project as being unique 

FY 2006 Progress Report 

and beneficial based on the project. The capability to 

develop a truck-quality Class A surface without lowprofile additives (which compromise toughness) and 

to achieve consistent molding is new to the industry. 

The excellent toughness seen in carbon hybrid composites is also an industry first, described with enthusiasm by the resin supplier as “ballistic 

performance.” Figure 11 illustrates the significant 

increase in impact energy compared with industry 

standards of sheet moulding compound (SMC) and 

liquid molding. One significant issue with carbon 

composites is their lack of toughness and brittle failure nature. The team had concerns about the minimum achievable panel thickness owing to this issue. 

Limiting panel thickness would severely hinder cost 

competitiveness but would provide desirable weight 

savings. Resin and process evaluation indicated a 

need for a system with close temperature control and 

required development of tight tolerances on the 

process and equipment. 

Impact Energy 

- first crack (J) 

3

2.5

2

1.5

1

0.5

0

Figure 11. Resin system allowing excellent impact energy. 

Filler Optimization in Hybrid Resin 

Filler optimization is important in minimizing 

viscosity for the lowest achievable fill pressures on 

large structures. Likewise, maximum filler loading is 

an important aspect of reducing resin shrinkage for 

achieving the best possible surface finish (Class A). 

The results of the DOE to optimize filler loading 

with viscosity included four factors: resin blend 

(20/80 to 30/70), filler load (25 pph to 45 pph). The 

filler ratio (15% to 35% filler A to filler B) meas145
FY 2006 Final Report 

ured at 2 and 12 min after mixing, as a function of 

viscosity, showed that there is an optimal region of 

filler ratio and filler loading within the tested ranges. 

In Figure 12, the Pareto chart of the effects, not including mixing time, shows resin blend as the largest 

determinant of viscosity at a given mix time. Therefore, the most sensitive component of final mix viscosity is the isocyanate-to-polyester ratios, with 

some dependence on filler load. 

High Strength Weight Reduction Materials 

step needed to verify and improve the resolution of 

the data is a full factorial concentrated around a focal point of a filler ratio of 23%; a filler load of 

35%; a resin blend minimizing isocyanate but controlled for its effects on the final composite properties; and mixing times of 2, 7, and 12 min. 

Model and Experimental Correlation 

A representative region of the specimens was 

used and discretized in 3-dimensional finite elements for the simulation of the tensile stress/strain 

responses. 

The panels were then laid out for specimen orientation and cutouts. Figure 13 shows the layout for 

the specimens. The specimens were oriented with 

one referenced edge of the panel and specimens 

were then removed; one set was parallel (0°) and the 

other was perpendicular (90°) to the referenced 

edge. 

Test couponswerecut from panels

Figure 12. A Pareto chart illustrating the combined and 

individual effects of the different constituents of the 

mixes. 

The other non-setting factors show narrow 

ranges of minimum viscosity. The filler load seems 

to have the second largest effect on mix viscosity, 

and in general, viscosity increases with increasing 

filler load. Filler load shows a decreased viscosity 

response between 30 and 40 pph and a resin blend 

ratio below where an interaction may be occurring 

with the filler ratio. 

Filler ratio, the percentage of filler A within the 

blend of filler A and filler B, gives a minimum viscosity within the range of 20% to 25% depending on 

the total filler load in the mix and set-time. This interaction resulted in a shift in the viscosity minimum 

from 25% to 20% in filler ratio as the filler load 

shifted from 30 to 40 pph. It is also affected by the 

setting of the resin over time by shifting the optimum filler ratio back from 20% to 25%. This same 

shift with set-time can be seen in the resin blend vs. 

filler ratio above. Is this shift an interaction between 

the surface of one of the fillers and a resin component? 

The approach to this screening study was intentionally broad within the area we now know as a 

more optimal range for the factors tested. The next 

at0 and 90 degree

90 degree directio

n 90 degree direction

Figure 13. Specimen layout on panels. 

Tensile and flexural specimen geometry followed the ASTM D638 Type I Tensile and ASTM 

D790 Four-Point Bending standards, respectively. 

Specimen profile measurements, including width 

and thickness, were taken and recorded for each 

specimen. The specimens were also asymmetric in 

the panel through-thickness. 

The double extensometers were then used to 

measure the axial elongation and the width strain 

during axial loading during tensile testing. From this 

double extensometer measurement, a 3-dimensional 

146
High Strength Weight Reduction Materials FY 2006 Progress Report 

strain response can be calculated. These data are 

then used to develop the elastic stiffness matrix and 

constitutive equations for model development. The 

flexural specimens were tested in 4-point bending. 

This allowed the modeler to be able to compare how 

well the prediction correlated with experimental 

data. 

Figures 14 and 15 present the simulated tensile 

stress/strain responses compared with the experimental results for the glass/vinyl ester and carbon/vinyl ester specimens tested until failure. The 

simulated curves were obtained using the elastic 

EMTA model (dashed line) and the incremental 

elastic-plastic EMTA model (solid line) implemented in ABAQUS. The glass/vinyl ester specimens could undergo higher deformations than the 

carbon/vinyl ester ones. In the former case (Figure 14), after linear behavior at small strains, the responses become nonlinear at higher applied stresses; 

whereas the non-linearity of the responses is negligible in the latter case (Fig. 15). The non-linearity 

results mainly from two different material origins. 

The first origin is damage by matrix cracking, fiber/matrix de-cohesion, and fiber pullout and breakage at the final stage. The second cause is the plastic 

deformations of the matrix material. In this report, 

damage was not modeled, and only the plasticity of 

the polymer matrix was accounted for. Figures 14 

and 15 show good correlations of the predicted results with the experimental values, except around the 

final loading stage in which excessive accumulations 

of damage have led to failure of the specimens. 

Flexural tests (based on the ASTM D790 standards) were conducted for glass/vinyl ester, carbon/vinyl ester, and carbon/glass/vinyl ester 

specimens. The latter are made of two layers: one 

carbon/vinyl ester and one glass/vinyl ester layer. In 

all specimens, the major fiber alignment axis is parallel to the width direction. Figures 16 and 17 show 

the predicted load/deflection curves compared with 

the experimental values also presented on the same 

figures for the glass/vinyl ester and carbon/vinyl ester specimens. Good correlations with the experimental results were obtained with the use of the 

Figure 14. Tensile stress/strain responses of the random 

glass/vinyl ester specimens. 

Figure 15. Tensile stress/strain responses of the random 

carbon/vinyl ester specimens. 

elastic-plastic EMTA model. In order to determine 

the effect of large displacements (which occur 

around the end of the loading) on the overall responses, we also carried out an elastic-plastic analysis using the large displacement (geometric nonlinearity) option in ABAQUS. The corresponding 

solution for the carbon/vinyl ester specimen is denoted as “NLG elastic-plastic” and is presented in 

Figure 17. Compared with the elastic-plastic solution of the same problem, it is noted that the geometric non-linearity is negligible. This is because the 

maximum deflection is of the same order of 

147
FY 2006 Final Report 

Figure 16. Flexural load/deflection responses of the 

random glass/vinyl ester specimens subjected to fourpoint bending. 

Figure 17. Flexural load/deflection responses of the 

random carbon/vinyl ester specimens subjected to 

four-point bending. 

magnitude as the specimen thickness. Consequently, 

it is not necessary to carry out elastic-plastic analyses accounting for geometric non-linearity for these 

specimens. 

Preform Development and Evaluation 

Several different preform styles have been considered and evaluated for the program to date. 

These include commercial and development variations to achieve the goals for fiber hybridization. 

High Strength Weight Reduction Materials 

Using the DOE/Automotive Composites Consortium P4 preforming cell at the NCC, hybrid flat 

preforms were made in a continuous process by 

modifying equipment to meet requirements. This 

was a distinctly different approach from the previous 

efforts where separate preforms of carbon and glass 

were made and later bonded together by NCC. Issues were seen in the distribution of binder in this 

approach, especially given the preferred flat tow 

form of carbon fiber being used. Molding of these 

panels indicated a lack of binder and/or a lack of 

fully developed strength for interlayer adhesion, especially in the glass fiber layer. Preforms were later 

developed for a multi-sided test tool, as well as the 

large test mold for full-scale evaluation trials. These 

confirmed the lack of development of binder 

strength at the interface, as the all-carbon preforms 

molded extremely well while the glass preforms 

showed issues with fiber washout and lack of preform integrity. 

The requirement for a discontinuous fiber preform that can accommodate multiple project objectives was a key issue in the project. Critical aspects 

included allowable loft and compressibility, capability to achieve a Class A surface, ease of fiber placement, ability to tailor for local additional carbon 

fiber reinforcement placement, minimum cost of 

processing, and repeatability of processing. The approach allows a given amount of loft to be available, 

yet is both compressible and able to provide a highpermeability resin flow path. The method and application of the surface fibers have provided benefits in 

tailoring fiber content and assisted in achieving success in the Class A arena. 

A preform fabrication cell was developed to allow rapid changes to the preform structure, and subsequent molding trials, to allow more efficient and 

cost-effective evaluation of specific approaches. 

Hybrid preforms were developed that provided 

complete mold filling and wet-out capability and 

provided samples for mechanical property testing. 

Work to make this process repeatable and predictable continues. 

Preforming development was completed on two 

fronts: (1) small test preforms for both the flat glass 

mold and multi-faceted test tools for determining 

permeability and resin/preform interactions and (2) 

large preforms for evaluation of scale-up factors and 

ability to meet design feature performance for subcomponent test tooling. 

148
High Strength Weight Reduction Materials 

To supply the various molding trial requirements, different core densities and fiber surface densities were evaluated. Packing levels were from net 

to 60% compaction, which varied fiber loading in 

the composite from effective values of 25% to well 

over 45%. Fiber loading in the carbon fraction was 

approximately 50% and allowed excellent surface 

density and fabrication of a thin, high-stiffness 

panel. This is clearly seen in the test results for flexural properties vs. the results for tensile properties. 

Figure 18 shows the surface available from the 

sample panel and several complex features that address draft angles, surface finish, and non-uniform 

thicknesses. The large tool evaluation will be explained in a separate section. 

FY 2006 Progress Report 

glass-top mold (Figure 19) was loaded with the preform and infused with both neat and filled blended 

resin systems using VARTM. The transparent mold 

allowed for visual confirmation of the location and 

flow path of the hybrid resin at any point during the 

infusion process. In addition, the mold was fitted 

with inlet and outlet pressure transducers to record 

pressure gradients during mold filling. The second 

mold was constructed and fitted to a hydraulic press 

operating at 2500°psi. Visual confirmation of resin 

flow was confirmed via short tube ports protruding 

from the upper mold half. The panel preforms were 

one of two types of cores, with and without chopped 

carbon fiber. In all cases, temperature was between 

40 and 50°C;and the mold cavity was evacuated to 

85 kPa prior to injection of the resin with an 11:1 

custom piston pump, which kept the unmixed resin 

components at 45°C. 

Figure 19. Glass top mold for test trials. 

Figure 18. Molded preform indicating 

discontinuous fiber surface. 

The alternative preform material showed significant improvement. Using combinations of the hybrid 

resin and the hybrid preforms, test panels were 

manufactured at different lengths to provide specimens and validate the feasibility of molding large 

(>5 ft) and thin (<3 mm) components. Close monitoring of developing manufacturing procedures provided valuable data concerning the behavior of both 

the resin and fiber hybrids in vacuum-assisted RTM 

(VARTM) and closed molding operations. 

Two types of flat laminate panels were produced 

from two different mold sets. Initially, a 12×24 in. 

Working with material manufacturers, four cores 

of different thickness and architecture were examined as possible core materials for use with the 

highly viscous blended resin. CR-1, FM-1, S&F and 

“M-03” core materials were compared during initial 

testing. Only the S&F products were capable of providing fully wet-out panels of less than 5 mm. S&F 

and M-03 are 4-mm and 3-mm- thick fabrics, respectively, consisting of glass strand mat stitched to 

a “knitted” polypropylene core, which provides loft 

to create a flow path for the resin and reduce backpressure within the mold. S&F designed these mats 

specifically for the purpose of liquid injection molding with high-viscosity, highly filled resin systems; 

and the fabric has never been tested previously, ex149
FY 2006 Final Report High Strength Weight Reduction Materials 

Scott & Fyfe Polymat High Flow cept during product development at S&F. Discontinuous carbon fibers were chopped from Toray 12-k 50 

tows directly onto the cores and held in place with 

polyester binder. In some cases, final placement and 

consolidation by hand was required to achieve full 

panel coverage and even fiber distribution. When Pressure (psig) 

30 

10 filler was introduced to the matrix, calcium carbonGate Pressure 

Injection 

Mold Full 

Vent Pressure 

Cure 

Mold under vac 

ate, in powder form, was added to the polyester side 100 110 120 130 140 150 160 170 180 190 200 210 220 230 240 250 260 270 280 290

of the blended resin. The resin itself was supplied as -10 

Time (s) 

a proprietary chemistry consisting of a polyester 

component, a compatible urethane component, and a Figure 20. Pressure difference between gate and vent 

catalyst. during injection cycle for S&F core material. 

The purpose of this investigation was to evaluate 

300 

Scott and Fyfe M-03 Glass Mat 

-10 

0 

10 

20 

30 

40 

50 

100 120 140 160 180 200 220 240 

Time (sec) 

Pressure (psig) 

Gate Pressure 

Vent Pressure 

the processability of the blended resin and hybrid fiber/mat raw materials and characterize the required 

production cycle. During the early stages of testing, 

the S&F mats performed exceptionally well, allowing the matrix to fill and fully wet the glass and hybrid preforms in less than half the time taken by a 

competitor’s core materials (Table 2). Figures 20 

and 21 show the difference between inlet and outlet 

pressures recorded during the RTM process. Both 

the S&F and M-03 had equal gate and vent pressures 

in less than 100 seconds, with the thinner M-03 requiring slightly more time. It is important to note the 

design of the mold cavity itself is critical to both the 

fill time and the level of filler filtering. Resin must 

be injected directly into the polypropylene core of 

the mat, along the material plane. Otherwise, a layer 

of glass strand hinders the flow rate within the S&F 

Figure 21. Pressure difference between gate and vent 

during injection cycle for M-03 core material. 

and it behaves similarly to any other core material. 

Table 2. Recorded fill times for core materials 

slightly along the inlet width, a problem that was 

corrected by pinching the fiber along its edge and 

cutting a small section away directly in front of the 

resin gate. 

The quality of the final panels was excellent in Material Time to fill 24×12× 0.12 in. cavity 

CR-1 4:38 

FM-1 5:20 

S&F 1:43 

M-03 0:34 

both the 2-ft and 6-ft panels. Panel surfaces showed 

no shrinkage or print–through, and thickness were as 

consistent as the molds would allow. Final panel 

thicknesses were between 2.5 and 3.25 mm when MThe success of the 6-ft panel validates the use of 

the filled, blended resin system and S&F M-03 for 

the production of large components. The injectiontook less than 4 min to complete, and the panel was 

demolded in less than 25 min from first injection. 

Resin ports along the length of the mold indicated an 

03 was used and between 3.5 and 4.5 mm when S&F 

was used. The nature of the S&F mat filling from 

“behind” the glass mat virtually eliminated washout 

of carbon fiber during RTM; however, polyester 

powder was still applied to ease preform handling. 

Filler improved the shrinkage characteristics 

slightly, although it would slow the rate of cure, 

possibly as a result of absorption of promoter over even flow rate along the full length of the panel. 

When demolded, the preform had been stretched time. The resin system also performed admirably, 

wetting and bonding to both types of fibers, as well 

as the polypropylene core. The fast fill rates and 

high level of promoter reduced cycle times to less 

than 16 min for the short panels and 25 min for the 

150 
High Strength Weight Reduction Materials FY 2006 Progress Report 

longer pieces, although the larger mold was frequently at a lower temperature than the glass-top. 

Panels produced from the S&F mat and carbon 

chop were sectioned and subjected to flexural and 

tensile testing. The results agree quite well with the 

predicted material properties for hybrid fibers and 

blended resins. In addition, optical analysis of panel 

cross sections showed no signs of filtering out of 

fillers from the resin along the length of the panels. 

Large Evaluation Tool 

In order to evaluate the capability of the process 

and materials to meet demands of industry, a prooftest component tool was developed. The tool, which 

has all the features anticipated in designs, was used 

to evaluate the hybrid perform process for achieving 

performance given the requirements of these features. It is not expected that all features will be 

molded and perform as expected, since the goal of 

the development is to determine the limitations of 

the hybrid fiber system and to provide designers and 

OEMs with guidelines for where lightweight materials will be applicable and how to design and manufacture in those cases. 

Project personnel ran the evaluation tool over a 

week-long period at a tier-one supplier, with support 

from the tier-one manufacturing floor personnel. 

Figure 22 illustrates the size of the test tool for 

evaluating complex preforms and mold filling. It 

was immediately apparent that this was not a desirable development process, as the systems for preforming, RTM, and resin optimization had to be 

developed independently and then brought in to the 

test. The P4 preforms developed at NCC showed 

great limitations. Alternative material systems were 

then evaluated that allowed the preforms to be modified. 

Results from the trials showed issues with the 

binder capability of the preforms. Interaction of the 

resin system softened the binder over time (this was 

not observed to such an extent in smaller-scale test 

tools, as the resin had much shorter contact times 

with the preforms before the smaller tools were 

completely filled and flow stopped). All-carbon fiber preforms were much more successful than glass 

fiber preforms, and several test sections were useful 

for fabrication of coupons, etc. Indications from the 

test were that resin temperature and tool temperature 

control were critical for this system, which led to 

Figure 22. Mold cell for test tool setup for 

large part molding. 

the resin development experiments described. Mold 

temperature was seen to be a significant issue; this 

finding has led to new approaches to tool design and 

construction that are being evaluated. 

A new cell was set up for R&D purposes to enable preforming and resin modification in a more 

appropriate environment. Also, several options for 

molding were built in that were not possible in the 

existing production environment but could easily be 

implemented once proven. In operating this cell, 

several successful test panel runs were made with 

hybrid preforms that met the requirements of the 

overall process. 

The newly developed hybrid preforms using 

commercial materials and adapting them to hybridization not only proved to be cost-effective for the 

OEMs but also significantly improved fill and cycle 

cure time. Several tests were run using large tooling 

and comparing different materials and flow behaviors. 

Based on the new material and test results, flow 

modeling was used for correlation in the large tool 

experiments. 

Flow Modeling 

Flow modeling used Polyworx software for 

simulations in the flat plate tools for determining 

preform permeability based on temperature, pressure, and viscosity. Figures 23a and 23b illustrate 

the pressure and flow patterns in the long, flat plate 

tool. The flow and pressure pattern was even. The 

151
FY 2006 Final Report 

Figure 23a. Modeled flow pattern in 18×72 in. panel 

mold. 

Figure 23b. Modeled pressure pattern in 18×72 in. 

panel mold. 

data obtained from the flat panels were correlated 

with experimental data, which provided the information needed for inputting into the more complicated 

test tool. 

The large, complex test tool was then modeled 

using the permeability numbers determined from the 

flat panels and the other resin input variables. The 

modeling of the test tool was used to show the correlation between the modeled and experimental results 

for scaling up complexity in the part. The simulations contributed to understanding of flow patterns 

and fill times with the current material configuration. This gave confidence that the modeling of the 

design will work prior to tool fabrication. 

High Strength Weight Reduction Materials 

Mechanical Testing and Analysis 

Testing and analysis were completed on all 

variations of material combinations. The project initially tested materials for the EMTA modeling that 

used the P4 preforms with varying levels of glass 

and carbon ratios with a low-cost vinyl ester resin 

system. The program then transitioned into the new 

hybrid resin materials and the hybrid preforms with 

the alternative performing approach. Tensile, flexural, and fatigue mechanical properties were tested 

for the different material combinations with some 

very satisfying results. 

The analysis used microscopy and Archimedes 

measurements for determining the void volume, 

filler filtering, and interpenetration into the fiber 

tows. The Archimedes measurements were used to 

determine density of the materials and variations 

across the large parts. Figure 24 illustrates the fiber 

wet-out and the filler interaction with the fiber tows. 

This image shows great penetration of the filler into 

the tows that gives a Class A surface finish. Sur- car face

-

corglas

Fiber 

48.12% 

Matrix 

49.93% 

Porosity 

1.88% 

Figure 24. Microscopy of hybrid material with filler and 

surface coating. 

The density of the hybrid resin and hybrid composite system used as the optimum setup was 25% 

less than that of SMCs used for semi-structural 

components and surface-finished parts and 15% less 

than the continuous strand mat (CSM) typically used 

in RTM of parts for OEMs. The tensile modulus was 

nearly 30% higher than that of the CSM. Flexural 

strength was approximately 50% higher than that of 

the CSM and equal to that of unsaturated polyester 

SMC. The flexural modulus, which is the key driver 

for designers in stiff driven parts, was three times 

that of CSM and 25% higher than that of SMC. Instrumented impact is over three times that absorbed 

by the unsaturated polyester SMC and compares 

152
High Strength Weight Reduction Materials FY 2006 Progress Report 

with that of a 50% glass loading SMC material that 

is nearly 25% heavier 

The other outstanding property was fatigue 

strength in the composite. The OEM typically does 

not go past 106

 cycles. The samples tested in the 

newly configured material were tested out to 107

 cycles at a 60% stress factor. This test exceeds all expectations of the material design. 

Tool Design Concept 

A tooling method for application in rapid fabrication of thin-walled composite structures was developed under this program to help reduce the high 

cost of steel tooling. 

During the investigation of layered hybrid composites, we investigated and proposed a low-cost, 

relatively rapid tooling system for RTM of polymer 

composites. The program evaluated the use of the 

MITTM (multiple insert tooling) system. We requested and received verbal authority from the 

North American technology licensor—JHM Technologies—to use MIT on the non-commercial development program. 

Our unique approach was to modify the MIT 

tooling and develop a MIT skin that achieved very 

high thermal conductivity, had responded rapidly to 

thermal excursions, had a near-zero CTE, and was 

stiffer than fiberglass skins. The rationale was that 

we needed to limit dimensional changes during the 

molding cycle and provide a system with increased 

durability in order to keep the surface of the tool in 

excellent condition over the tool life (or extended 

tool life), as well as provide for temperature control 

of the specific resin system chosen for the molding. 

A tooling material approach was used in which 

the traditional thin molding surface skin (A-side or 

appearance mold skin) was fabricated with near-zero 

CTE and a high thermal conductivity. This skin replaces the typical skin of fiberglass composite in the 

MIT system (Figure 25). 

The MIT skin uses a bolster to capture an independent skin that acts as a movable mold surface. 

This allows the operations associated with skin and 

fiber preparation to be done independently of the 

primary tool and press structure. Skins can be easily 

and cheaply fabricated from a pattern master, and 

multiple skins can be used to optimize the process, 

matching the press cycle to the number of skins required for the preparation cycle. 

Figure 25. Multiple insert tooling approach using highly 

thermally conductive skins. 

The approach of using a carbon MIT skin is 

novel in several ways. First, it provides a skin surface that is dimensionally similar at all temperatures. 

This ensures that the skin dimensions are similar 

during application of mold coat, fiber perform, installation in bolster, and molding and cool-down cycle. Internal stresses and surface dimensional 

changes are eliminated. Second, it provides for very 

rapid heat-up and cool-down of the mold cavity as 

heat is transferred very efficiently through the skins 

and into the part. This works in conjunction with the 

hybrid fiber and other fiber molding approaches to 

ensure consistent and repeatable process cycles and 

rapid processing capability. Third, the skins are 5–8 

times stiffer than a typical glass fiber skin because of 

the improved fiber stiffness. Large, thin skins are 

more easily handled outside of the bolster, and their 

lower deflection leads to improved durability and 

less likelihood of surface cracking of the tooling surface coat. This is a major factor in mold surface 

quality deterioration and one that helps to resolve a 

related factor: deflection over time leads to “patterning through” in glass fiber molds, a process in which 

repeated flexing of the mold skin causes fibers near 

the surface to be forced against the surface and begin 

to appear in profile. This is eventually passed into 

the part. 

Furthermore, high-temperature resin systems 

that have been modified offer better thermal conductivity and more uniform heating across the skin sur153
FY 2006 Final Report 

face for improved part process and controls. The 

high-temperature resin systems also provide higher 

heat deflection temperatures and resistance to preform fiber surface damage to the tool while the 

resin’s exothermic reaction is taking place. The 

high-temperature resin system must be able to handle exothermic temperatures greater than 150°C. 

The concept was also extended to the other 

components of the system, specifically the B-side 

tool and the bolster. Higher and more effective heat 

transfer rates and process lags of fewer than 3°C 

were noted, which improved the quality of molding 

significantly. A notable benefit of this approach is 

the capability to achieve consistent cure times and 

eliminate local hot spots caused by varying thicknesses of laminates. The process benefits translate 

into limiting pre-release from the tool surface and 

achieving better dimensional control and surface 

quality. 

A set of durability tests was performed on some 

laminated test pieces with a surface coating that will 

allow us to achieve a Class A or near–Class A finish 

on the surface. The series included testing three 

specimens to failure in 4-point bending. This determined the first ply failure of the laminate and the 

load at which the surface coating cracked. Testing 

was then performed on the laminate in cyclic fatigue 

using the same 4-point bending setup and setting the 

peak load at 30%, 40%, and 50% of the first ply 

failure average of the previous testing to failure. 

These samples were then subjected to a 3-Hz cyclic 

fatigue. The samples reached 500,000 at 30%, 

200,000 at 40%, and just under 100,000 at 50% 

when the test was terminated because surface cracking was observed. The durability of the carbon fiber 

laminate was very good, and there was little loss in 

the load for the set deflection. 

Large Tool Scale-up 

Two approaches were taken for tooling on a 

hood design. The first was an MIT approach that allowed for offline surface prep and de-molding separately from the press, which helped keep cycle times 

low and throughput high. The second approach was 

VARTM using lighter and lower-cost tooling to accomplish the same result while using the thermally 

conductive skins developed for the MIT approach. 

The difference in cost between these two tools 

was nearly nine times, with VARTM being the 

lower-cost method. The challenge in doing very 

High Strength Weight Reduction Materials 

large structures is the weight of the tooling and the 

company’s ability to handle such large, heavy tooling. Figure 26 illustrates the hood of a heavy truck 

that was planned for testing. 

Figure 26. Proposed hood for tooling cost comparison. 

The project had started building the hood tool 

with a tier-1 supplier to the OEM partner. The company went into receivership and the OEM’s competitor took over the company. The tooling was 

removed and held in storage while the competitor 

was working on selling the company. The new 

owner of the company re-quoted the tooling, and the 

price was much higher than the initial tooling quote. 

The inflated cost eliminated any work with the new 

owners of the company. Figure 27 shows the progress that was made up to the point of receivership. 

Figure 27. Master pattern for tool builds 

The OEM tried several other tier-1 suppliers in 

its supply chain with little success. The supply chain 

154
High Strength Weight Reduction Materials 

is extremely tight at the current production capacity 

in the trucking industry, and there is no shop time 

available to take on projects. The cooperative research and development agreement (CRADA) partner preferred to have a tier-1 supplier do the work so 

the partner could incorporate the technology into its 

product lines and be trained with the new technology. 

An alternative supplier was found and a reasonable cost estimate given for demonstrating the technology with VARTM and our new skin technology. 

However, the CRADA partner management did not 

wish to engage any suppliers outside the preferred 

supplier base. 

The current estimates for cycle times on a 72-ft2 

part are near 15–20 min. Estimated fill times are 4– 

9 min, depending on the volumetric flow rates of 

the pump. These cycle times compare with nearly 3 

hours on a laminated hood and about 15 min on a 

preform hood (in steel molds at a cost factor of almost 5×). 

Conclusions 

The flow trials gave us valuable insight as to 

how fast the large panels will fill. The rheology 

studies indicate time sensitivity when the two resin 

components are mixed and have the largest influence 

on the viscosity. There is an optimum ratio of the 

two fillers at 25 pph that will help in reducing filler 

cost and viscosity. The viscosity data became very 

useful with the modeling, which illustrated the fill 

time and pressure gradients within the part. These 

models and the correlation data gave us the information needed to model large structures prior to tool 

fabrication, which helps reduce the risk of the unknown for large, complicated, and expensive tools. 

Future work will model the large composite 

structure and build the tooling. An experimental correlation will be evaluated for the large molded parts. 

A large tool will be fabricated for testing the moldability and mechanical testing of the large parts. A 

series of factors associated with optimizing details of 

the process technology are being analyzed, and the 

CRADA partner is developing an approach to mold 

full-scale parts that will ultimately go on over-theroad trials. We chose the parts to demonstrate the 

most challenging aspects of the process capability, 

namely structural Class A surfaces in direct line-ofsight with some complex process details required for 

success. Tooling quotes and negotiations are ongoFY 2006 Progress Report 

ing; we anticipate that several aspects of the process 

will go forward for patent protection. 

The program has generated a wealth of information. The goal of reducing weight reached a reduction of 40% of a replacement component 

comparison. The tooling cost competed with that for 

a single-sided chopper gun tool, which is several orders of magnitude lower than the steel tooling for 

SMC material processes. The cost of carbon fiber 

continues to be a serious problem for the adoption of 

new technologies. The cost of carbon fiber at the beginning of the program was $6.50/lb and climbed as 

high as $14/lb. Calculations for one part indicated 

that a break-even point for this new technology 

would be $6/lb. If the material price would return to 

reasonable levels, the adoption of new, lightweight 

hybrid systems would be an attractive investment. 

Project accomplishments included 

• predicting properties in chopped fiber composite 

RTM preform moldings using the newly 

developed EMTA modeling code 

• modeling flow behaviors over large structures 

• developing a unique approach to preform 

fabrication for increased fill speeds for large 

structures 

• developing a hybrid resin for rapid cycle times 

• developing new tooling skins for thermal control 

of exotherms, and class A surface finish that 

have combined to create increases in the 

mechanical properties 

A computational tool using the EMTA approach 

was developed for thermo-elastic and elastic-plastic 

analyses of short-fiber polymer composites. It was 

created by implementing the EMTA-based constitutive models into ABAQUS. The simulations of tensile and flexural tests conducted on glass/vinyl ester, 

carbon/vinyl ester and carbon/glass/vinyl ester 

specimens provided good predictions of composite 

response and were in agreement with the experimental results. The computational tool can be used to determine the basic properties needed for the design of 

short-fiber polymer composites, such as effective 

thermal and elastic properties and tensile and flexural moduli of the composite laminate. Moreover, 

the ability to account for plastic deformations using 

the incremental elastic-plastic EMT model enables 

modeling of the composite behavior at high loading 

levels leading to deformations that exceed the elastic 

limit. The model inputs are fiber (or fiber tow) and 

155
FY 2006 Final Report 

matrix properties, fiber (or fiber tow) volume fractions, aspect ratios, and the orientation parameter 

(θ). This computational tool offers the flexibility to 

adjust these parameters so that the resulting homogenized properties and responses satisfy the design criteria in terms of desired properties and 

deformation limits. The tool can therefore be used to 

assist in the design and manufacture of short-fiber 

composite structures. 

The preform development started with a preforming process that would not work on large structures because of resin flow resistance. Development 

of a new preform system that used commercial materials was successful and helped achieve relative fill 

times nearly 10 times faster than those of previously 

used preform technologies. 

The new hybrid resin system developed in the 

lab significantly improved the fiber wetting and 

bonding of two dissimilar fibrous materials. This 

helped improve mechanical properties by as much as 

50% and increased the impact strength by three 

times compared with commercial materials. 

Development of new tooling demonstrated that 

maintaining temperature control is extremely important, as the material exothermed during curing. Heat 

buildup in the tooling limits the fill time for the tool 

because of resin rheology changes. Tooling fatigue 

High Strength Weight Reduction Materials 

life was another important consideration: as large 

tools are handled and flexed, the fatigue can cause 

surface crazing and spider webbing that will impact 

the molded surface in the new part. 

The OEM considered this program a technical 

success, but issues with tier-one suppliers and capacity limitations created by current high truck sales 

limited the transfer of technology to a supplier. The 

OEM is considering this material for future use in 

smaller parts as a result of the information gained 

from this program. 

Presentations/Publications/Patents 

B. N. Nguyen, K. L Simmons, G. J. Grant, and 

M. T. Smith, “A Computational Tool for Analysis 

and Design of Short-Fiber Polymer Composites,” 

prepared for submission to Composites Science and 

Technology, 2003. 

ASM International, Engineered Materials 

Handbook: Composites, ASM International, Metals 

Park, Ohio, 1987. 

Three invention disclosures were submitted under this program. 

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High Strength Weight Reduction Materials FY 2006 Progress Report 

B. Hybrid Composite Materials for Weight-Critical Structures 

Principal Investigator: Curt A. Lavender 

Pacific Northwest National Laboratory 

P.O. Box 999 Richland, Washington 99352 

(509) 372-6770; fax: (509) 375-4448; e-mail: curt.lavender@pnl.gov 

Principal Investigators: Donald Trettin 

PACCAR Technical Center 

12479 Farm to Market Road, Mount Vernon, Washington 98273 

(360) 757-5241; fax: (360) 757-5370; e-mail: don.trettin@paccar.com 

Chief Scientist: James J. Eberhardt 

(202) 586-9837; fax: (202) 587-2476; e-mail: james.eberhardt@ee.doe.gov 

Field Technical Manager: Philip S. Sklad 

(865) 574-5069; fax: (865) 574-6098; e-mail: skladps@ornl.gov 

Contractor: Pacific Northwest National Laboratory 

Contract No.: DE-AC06-76RLO1830 

Objective 

• Develop and demonstrate (1) the application of hybrid composites and composite/metal hybrids to heavy-duty 

vehicles and (2) the capability to integrate these materials choices into moderate-volume production. 

• Develop and demonstrate the potential for major weight savings (>50% on a component basis) in critical 

structures applicable to truck cabs and support components. 

• Demonstrate the basis for the use of hybrid metal-composite systems to reduce weight via proof-of-principle 

experimentation. 

• Develop full-scale prototype components for vehicle testing and validation. 

Approach 

• Investigate the potential of new materials and manufacturing technologies to effect major weight reductions for 

heavy-duty vehicles. 

• Assist in demonstrating the applicability of composites and composite/metal hybrids to operational vehicles 

with little or no cost impact. 

• Provide the experience base to develop design and analysis tools, as well as the scientific understanding of the 

factors affecting molding and materials performance. 

• Provide the materials suppliers with a market that can stimulate demand, leading to an increase in their 

production capacity. This will help reduce materials costs by creating higher volumes. 

• Develop and assemble a full-scale prototype door system using hybrid material for heavy vehicle testing and 

evaluation. 

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FY 2006 Progress Report High-Strength Weight Reduction Materials 

Accomplishments 

• Provided technical input to PACCAR for further implementation of the “hybrid” materials technology into a 

production vehicle. 

• Provided technical information for the development of a PACCAR supply chain for the hybrid materials 

approach. 

• Completed three sets of prototype hybrid door components, which were delivered to Pacific Northwest 

National Laboratory (PNNL) by the subcontractors. The key components include sand cast magnesium door 

inners (Style and Tech) and molded hybrid glass/carbon fiber upper door frames (Profile Composites). 

• Developed bonding fixture, bonded the initial hybrid door components, and verified dimensional conformance. 

• Assembled the first door and performed static load analysis. Initial results indicate that the hybrid door 

assembly exceeded the targeted 50% improvement and exhibited a 64% improvement. 

• Assembled additional door for cab fit and function and full-length cab door durability testing. 

Future Direction 

• Support implementation of the hybrid materials system in a PACCAR production vehicle. 

• Compile a final report at the conclusion of the prototype demonstration phase. 

Introduction 

Current materials and manufacturing technologies used for heavy vehicle door systems are often 

dictated by the high cost of tooling and the relatively 

low production volumes for Class 8 trucks. Automotive-style stamped door designs, whether of steel or 

aluminum, require multistage stamping dies that are 

generally cost-prohibitive at lower production volumes (<50,000 units per year). Alternate materials, 

such as glass-reinforced sheet moulding compound 

(SMC), require less expensive tooling and can provide a Class A finish, but the relatively poor specific 

properties of SMC tend to compromise design and 

result in a heavier door system. For many production 

truck cabs, a simple aluminum extrusion frame is 

used with a flat aluminum sheet riveted to the frame. 

Although this approach does not require expensive 

tooling, the use of constant cross-section extrusions 

in the frame is less than optimum, and it requires 

more assembly labor than other approaches. PACCAR, a world leader in Class 8 truck design and 

manufacturing, teamed with PNNL to explore alternate “hybrid” door system designs that minimize 

tooling cost and per/part door cost, while providing 

a lightweight, structurally stiff, automotive-style 

door. 

Project Approach 

The initial approach to development of the hybrid door system was to perform a structural analysis 

of an existing PACCAR door design and to determine what the design and performance goals should 

be for new-generation door systems. PACCAR provided a number of weight, cost, and performance 

parameters that it considered important for future 

door designs. PNNL was tasked with surveying existing and emerging materials and manufacturing 

approaches that could be applied to a new door design. Following completion of this survey and 

analysis of existing door designs, PNNL, with design assistance from Mercia, Ltd., developed a series 

of five door design concepts that included combinations of large die castings, extrusions, carbon- and 

glass-reinforced composites, and conventional SMC 

and stamped aluminum exterior panels. 

Following a concept review meeting with PACCAR, an optimized hybrid door design concept was 

selected. The door concept was then defined using 

computer-aided-design tools and analyzed with finite element models to validate performance, 

weight, and cost. After determining that the prototype design met or exceeded all performance and 

projected cost targets, PNNL and PACCAR selected 

methods to produce prototype components for the 

full-scale assembly and testing phase of the project. 

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High-Strength Weight Reduction Materials FY 2006 Progress Report 

The finite element model of the prototype door system is shown in Figure 1. 

Figure 1. Finite element model of prototype door system under simulated 

loading conditions. 

Following completion of the design selection 

phase of the project, full-scale hardware components 

were developed and shipped to PACCAR Technical 

Center for assembly and testing. In addition to conducting the static deflection tests, PACCAR assembled a complete door and mounted it in a Class 8 

truck cab for fit and functional tests (Figure 2). Subsequently, the door was placed into a door durability 

cab test, which is currently being run to completion. 

Conclusion 

The team and its selected subcontractors completed the prototype development of hybrid door 

components for an advanced heavy truck door. The 

team has completed prototype assembly, including 

inspection and assembly fitting, as well as adhesive 

bonding development. The team assembled three 

prototype doors for cab testing and evaluation. Static 

testing and cab functional tests and evaluation have 

been completed. A prototype door has successfully 

completed durability testing, passing all requirements while reducing the mass by 37%. 

Figure 2. Prototype door assembled in Class 8 

cab for functional testing. 

The use of the hybrid materials approach is under evaluation by PACCAR for use in future vehicle 

systems. When required, PNNL has 

provided additional analysis and data to support implementation. 

159
High Strength Weight Reduction Materials FY 2003 Progress Report 
High Strength Weight Reduction Materials FY 2006 Progress Report 

C. High-Conductivity Carbon Foam for Thermal Management in 

 Heavy Vehicles 

Principal Investigator: B. L. Armstrong 

Oak Ridge National Laboratory 

P.O. Box 2008, Oak Ridge, TN 37831-6063 

(865) 241-5862; fax: (865) 574-6918; e-mail: armstrongbl@ornl.gov 

Principal Investigator: A. D. McMillan 

Oak Ridge National Laboratory 

P.O. Box 2008, Oak Ridge, TN 37831-6085 

(865) 241-4554; fax: (865) 576-0003; e-mail: mcmillanad@ornl.gov 

Chief Scientist: James J. Eberhardt 

(202) 586-9837; fax: (202) 587-2476; e-mail: james.eberhardt@ee.doe.gov 

Field Technical Manager: Philip S. Sklad 

(865) 574-5069; fax: (865) 574-6098; e-mail: skladps@ornl.gov 

Contractor: Oak Ridge National Laboratory 

Contract No.: DE-AC05-00OR22725 

Objective 

• Develop an understanding of the demands for heat rejection for heavy vehicles and develop designs that utilize 

the superior performance of high-thermal-conductivity graphite foams. 

• Coordinate efforts with an engine manufacturer so as to target the design(s) toward their concepts of the heat 

loads for heavy vehicles. 

Approach 

• Study fundamental mechanisms of heat transfer using graphite foam. 

• Develop a testing method to evaluate the foams and designs using the foams. 

• Work with an industrial partner to develop complete understanding of heavy-vehicle-related issues. 

Accomplishments 

• Achieved the production of graphite foam with homogeneous, controlled structure, allowing for tailorability for 

a given application, broadening the applicability of graphite foam in the marketplace. 

• Developed a heat transfer model. It was determined that pressure drop across the foam is proportional to the 

flow length of the foam. In addition, permeability (and thereby, structure) of the foam plays a role in the 

thermal performance. 

Future Direction 

• Evaluate potential on-vehicle tests to be performed at industrial partner’s facilities. 

161 
FY 2006 Progress Report High Strength Weight Reduction Materials 

Introduction 

A novel technique for creating pitch-based 

graphite foam was developed at Oak Ridge 

National Laboratory (ORNL). This technique uses 

mesophase pitch as a starting material but does not 

require the costly blowing or stabilization steps 

seen with typical carbon foams. 

The ORNL foam is an open-cell structure with 

highly aligned graphitic ligaments; studies have 

shown the typical interlayer spacing (d002) to be 

0.3356 nm, very near that of perfect graphite 

(0.3354 nm). As a result of its near-perfect 

structure, thermal conductivities along the 

ligament are calculated to be approximately 1700 

W/m•K, with bulk conductivities ≥180 W/m•K. 

Furthermore, the material exhibits low densities 

3

(0.25–0.6 g/cm ) so that the specific thermal 

conductivity is approximately four to five times 

greater than that of copper. The very high surface 

2 3

area (20,000 m /m ) combined with the high 

thermal conductivity suggests that graphite foam 

has significant potential for application as a 

thermal management material. 

One issue of importance with regard to the 

performance of the ORNL foam is homogeneity of 

density. Previous tests have shown that the density 

gradient usually observed in both experimental 

and commercial foams adversely affects 

performance. Efforts were undertaken to reduce or 

eliminate the gradient from ORNL foams so that a 

thermal model can aid in the design of a heat 

exchanger device using the graphite foam as well 

as predict the effectiveness of the device. 

Experimental 

Design and control of the structure of highthermal-conductivity-graphite foam has been a 

technical hurdle since its development in the early 

1990s. For some uses, it is desirable to have a 

more open structure (more and/or larger 

interconnected pores) while maintaining some 

mechanical strength and high specific thermal 

conductivity. By modifying processing 

parameters, such as foaming pressures and/or 

inorganic additives to the precursor material(s), 

such structure control can be realized, i.e., specific 

processing parameters can be tailored to specific 

performance properties (e.g., pressure drop, 

permeability), resulting in a high-thermalconductivity graphite foam that is potentially less 

friable and has controllable pore and pore window 

sizes. 

ORNL produced a volume of graphite foam 

pieces based on the needs of our CRADA 

partner’s application. The foam material was 

processed in accordance with an ORNL 

proprietary process. This material was shipped to 

our partner’s subcontractor, ThermalCentric, for 

evaluation and fabrication of a prototype heat 

exchanger. A combined total of 1524 in.3

 was 

delivered to the CRADA partner and 

ThermalCentric. 

Subsequently, ThermalCentric used the 

ORNL-produced graphite foam to assemble a 

prototype heat exchanger. The design of the heat 

exchanger was in response to performance 

parameters supplied by our CRADA partner. This 

device was subjected to the following test 

parameters in an air-to-air heat exchange test at 

our partner’s facility. Hot-side inlet air flow was 

supplied at approximately 400°C and cool-side 

inlet air flow at approximately 30°C. The inlet 

temperatures and cold flow were held constant 

while the hot flow was varied through 50, 75, 125, 

and 150% of the design condition flow rate. This 

procedure was repeated, holding the hot flow 

constant and sweeping the cold flow through the 

same percentage range variation. Additional data 

were collected at the design condition at 200°C 

and 300°C.Results showed the hot-side inlet flow 

at 400°C was cooled to 37°C with a cold-side inlet 

flow of 25°C. Measurements of design-point 

performance were obtained at the beginning and 

end of the test program, and both results were 

identical. This demonstrates the heat exchanger 

was not damaged by internal thermal stresses 

despite the large temperature difference applied to 

its elements. 

Conclusions 

A variety of processing variables were 

evaluated in order to fabricate foam with 

controlled pore sizes and pore windows. A wide 

range of pore sizes and morphologies were made 

through control of these variables. Examples of 

the range of the pores size and morphology that 

are feasible through processing control are shown 

162 
High Strength Weight Reduction Materials 

in Figures 1–3. The micrographs taken on a 

scanning electron microscope (SEM), shown in 

Figures 1–3, represent three foams made at 

ORNL. A summary of their corresponding 

properties is provided in Table 1. A commercially 

available foam (Poco) is shown in Figure 4. Note 

the difference between the pore sizes in Figures 1– 

3 and the pore sizes in the Poco foam in Figure 4. 

Figure 1 (ORNL A) has the largest pore size but 

also the most uniform size across the thickness of 

the foam (not easily seen in this magnification). 

The pore windows of ORNL A are approximately 

the same size as the Poco pore size. Not only is the 

pore window size equivalent to that of the Poco 

foam in ORNL B, but the quantity of pore 

windows is also substantial higher, explaining the 

higher permeability seen in this foam structure. A 

biomodal distribution of pore size becomes more 

evident in the ORNL foams B and C, although it is 

not as evident in ORNL B. Also of interest in all 

the ORNL foam structures is the apparent absence 

of closed pores. The reduction of closed pores and 

the increase in pore window size and quantity 

directly relate to the increase in structure 

permeability. The average pore size and bulk and 

calculated thermal conductivity properties were 

measured at ORNL, and the permeability and 

Forchheimer coefficient were measured at 

ThermalCentric. 

The four foams, ORNL A, B, and C and Poco, 

were tested at ThermalCentric using its thermal 

transfer rig. Testing details have not yet been 

disclosed by ThermalCentric. The results, as 

reported orally by Brian Thompson of 

ThermalCentric, showed that even in the slightly 

bimodal structure of ORNL B, approximately a 

200% improvement over any other graphite foam 

was realized. This comparison is based on data 

from commercial vendors as well as those reported 

in the open literature. ORNL A had even more of 

an improvement in performance relative to ORNL 

B; however, the thin walls of the pores due to 

increased pore size and windows weakened the 

compressive strength of the foam structure so that 

it would not be useful according to the prototype 

design specifications. 

The prototype design specifications were 

supplied by the partner to demonstrate a proof-ofFY 2006 Progress Report 

Figure 1. Scanning electron micrograph of ORNL 

foam A. 

Figure 2. Scanning electron micrograph of ORNL 

foam B. 

concept application of foam. As a result, the 

partner commissioned a 1⁄8-scale heat exchanger 

prototype from ThermalCentric. The prototype 

was designed with a predicted effectiveness of 

80% and achieved an effectiveness of ~94% at the 

design point. The repeatability of the points was 

very good. The pressure drop was also lower than 

modeled. The reasons for this may include better 

163 
FY 2006 Progress Report High Strength Weight Reduction Materials 

Figure 3. Scanning electron micrograph of ORNL 

foam C. 

Table 1. Properties of ORNL foams made using 

varying processing variables 

Foam Foam Foam Property ORNL-A ORNL-B ORNL-C 

Avg. pore size 1131 714 978.5 (μm) 

Bulk 

calculated 

thermal 

conductivity 19.7+/-1.8 34.7+/-2.9 36.5+/-2.2 

(W/m K) 

X direction 

Bulk 

calculated 

thermal 

conductivity 16.4+/-0.4 33.0+/-1.3 35.8+/-0.8 

(W/m K) 

Y direction 

Bulk 

calculated 

thermal 

conductivity 26.8+/-1.2 79.1+/11.7 62.6+/-2.9 

(W/m K) 

Z direction 

Permeability Not 

(K m2

) 2.05E-09 6.73E-10 reported 

Forchheimer Not coefficient .3050 .8249 reported cf 

Figure 4. Scanning electron micrograph of commercial 

Poco foam. 

convective heat transfer, higher fin efficiencies, 

advantageous longitudinal conduction through the 

foam, and lower contact resistance in the larger 

prototype than in smaller test coupons. 

ORNL and its CRADA partner have 

successfully built and tested a graphite foam airto-air heat exchanger. The heat exchanger 

transfers more heat than predicted, easily survived 

the rigorous test procedures, and demonstrated 

very good repeatability of the design point over 

time. The model underpredicts performance and 

pressure drop. 

Presentations/Publications/Patents 

“Controlling the Structure of Graphite Foam 

Using Secondary Inorganic Additives,” ORNL 

invention disclosure. 

“Design and Control of Structure of Graphite 

Foam,” ORNL invention disclosure. 

164 
High Strength Weight Reduction Materials FY 2006 Progress Report 

APPENDIX A: ACRONYMS AND ABBREVIATIONS 

2-D two-dimensional 

3-D three-dimensinoal 

A aging 

AES auger electron spectroscopy 

AJT Advanced Joining Technologies, Inc. 

Al aluminum 

ANL Argonne National Laboratory 

APS Advanced Photon Source 

ASC advanced squeeze casting 

CAE computer-aided engineering 

CDM continuum damage mechanics 

CPM cycles per minute 

CR cold rolling 

CRADA Cooperative Research and Development Agreement 

CSM continuous strand mat 

CTEs coefficients of thermal expansion 

CT Scans computed tomography scans 

DARPA Defense Advanced Research Projects Agency 

DBSD electron back-scattered diffraction 

DIC digital imaging correlation 

DOE Department of Energy, design of experiments 

DTA differential thermal analysis 

EBSD electron back-scatter diffraction 

EDM electrical discharge machining 

EDX energy-dispersive X-ray 

EMTA Eshelby/Mori-Tanaka approach 

EPS expanded polystyrene 

FEA finite element analysis 

FE-SEM field emission scanning electron microscope 

FIB focused ion bean 

FSJ friction stir joining 

FSJ/P friction stir joining and processing 

FSP friction stir processing 

FSSW friction stir spot welding 

GBS grain-boundary sliding 

HAZ heat affected zone 

HDI high-density infrared 

HSWR High-Strength Weight Reduction 

165 
FY 2006 Progress Report High Strength Weight Reduction Materials 

ICP-AES inductively coupled plasma–atomic emission spectroscopy 

ICSAM International Conference on Superplasticity in Advanced 

Materials 

INEEL Idaho National Engineering and Environmental Laboratory 

IPM inches per minute 

IR infrared 

ISSR interferometric strain/slope rosette 

ITP International Titanium Powder 

LANL Los Alamos National Laboratory 

LDH limited dome height 

LPPM low-pressure permanent mold 

Mg magnesium 

Mg17Al12 brittle inter-metallic phase 

MgO magnesium oxide 

Mg2Si magnesium silicide 

MMC metal matrix composite 

Mo molybdenum 

NCC National Composites Center 

NGF next-generation frame 

NVH noise, vibration, harshness 

OEM original equipment manufacturer 

OFCVT Office of FreedomCAR and Vehicle Technologies 

ORNL Oak Ridge National Laboratory 

PFA progressive failure analysis 

PIVAC pressure infiltrated vacuum assisted casting 

PLI pressureless infiltration 

PNNL Pacific Northwest National Laboratory 

PTC PACCAR Technical Center 

PU/SUV pickup/sport utility vehicle 

R & D research and development 

RPM rotations per minute 

RSST reversed shoulder spiral tool 

RSW resistance spot welding 

RTM resin transfer mold 

RWMA Resistance Welder Manufacturers Association 

SAE standard American equivalent 

SC squeeze cast 

SDSMT South Dakota School of Mines and Technology 

SEM scanning electron microscope 

SMC sheet moulding compound 

SPF superplastic forming 

SPH smooth particle hydrodynamics 

SPR self-piercing rivets 

SrTiO3 strontium titanate 

166 
High Strength Weight Reduction Materials FY 2006 Progress Report 

SSBT subscale brake testing 

ST solution treating 

TEM transmission electron microscopy 

TTU Tennessee Technological University 

Ti titanium 

Ti-6Al-4V popular alloy of Ti that contains additions of 6 wt % aluminum and 

4 wt % vanadium 

TWB tailor welded blank 

UW ultrasonic welding 

VAR vacuum arc remelting 

VARTM vacuum-assisted resin transfer molding 

VHP vacuum hot pressing 

WTI Walford Technologies, Inc. 

167 
FY 2006 Progress Report High Strength Weight Reduction Materials 

168 

A Strong Energy Portfolio for a Strong America 

Energy efficiency and clean, renewable energy will mean a stronger economy, a cleaner environment, 


and greater energy independence for America. Working with a wide array of state, community, industry, and 


university partners, the U.S. Department of Energy’s Office of Energy Efficiency and Renewable Energy 


invests in a diverse portfolio of energy technologies.


For more information contact:


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1-877-EERE-INF (1-877-337-3463)


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