Question

Assume that you are appointed by a university as the technical manager of the University's Formula-SAE-A Race Car project and you are required to:

a) Identify the relationships between operational objectives and functional requirements for the case of a new Formula-SAE-A Race Car (F-SAE-A) in 2018. Cite three operational objectives and the functional requirements that are needed to realise these objectives via using qualitative measures only; and

b) What is meant by “MOE”? For the effectiveness analysis of the F-SAE-A, list what you think would be the 10 most important characteristics that should be exercised and measured in the analysis. For one of the MOE of the F-SAE-A, describe an operational scenario for obtaining a measure of its effectiveness.

Answer 1

housings. Indeed, carbon/epoxy composites possess the

superior combination of high specific strength, almost

limitless formability, and specific stiffness required to handle

the extreme stresses resulting from accelerations on the order

of four times the force of gravity while conforming to a

contorted aerodynamic shape and maintaining precise

geometric alignment at all times.

Even though composites have become ubiquitous in high-

performance race car structures, most companies and teams

keep their construction techniques proprietary. Most available

information concerning the design, construction, and repair of

composite structures is published by the aerospace industry,

which deals with substantially different geometries as well as

loading and safety requirements.

Over the past few years, the USNA FSAE team composite

design efforts have ranged from simple fiberglass non-

structural fairings to detailed studies on carbon fiber

suspension components [5]. As a logical progression, the

USNA FSAE team elected to go forward with a composite

monocoque frame for the 2011 and subsequent competitions.

This paper serves to document the design and fabrication of

the monocoque frame. In the interest of brevity, most of the

validation tests and results have been omitted in this paper

but are well documented for the interested reader to

investigate [6, 7]. It is the authors' hope that other teams with

limited composites design and manufacturing experience can

enjoy the vehicle performance benefits made possible with a

carbon fiber reinforced plastic (CFRP) monocoque frame.

DESIGN PROCEDURES

This section outlines the design procedures followed in the

construction of the composite monocoque frame. First, the

overall monocoque type and dimensions were determined.

The design process continued with understanding the

applicable equivalency requirements needed to safely replace

a traditional steel space frame with a CFRP monocoque

frame. Next composite materials and layup techniques were

selected based on advertised material properties and

preliminary analysis tools. Once design choices were

validated using various component level tests, the monocoque

was completed.

Monocoque Type

For this class of vehicles, there are various ways to

incorporate composite materials into the design. The simplest

method is to replace some of the steel members of a

traditional space frame with composite shear panels on the

sides, top and bottom of the vehicle. These panels provide a

modest weight reduction while adding some torsional rigidity

to the frame. The other extreme is to replace nearly all of the

steel tubing with a full monocoque frame. The competition

rules still require a steel roll hoop for driver protection so it is

not possible to replace all of the steel tubing. However, this

approaches would likely result in the greatest weight

reduction and also the greatest increase in rigidity.

Unfortunately this method is also the most complex and

requires extensive tooling. A more moderate approach is to

mate a steel rear space frame module to a carbon monocoque

forebody. This hybrid approach should result in significant

weight savings and an increase in torsional rigidity over the

steel space frame. Figure 1 shows the relative benefits

associated with these various approaches. The Naval

Academy team elected to use the hybrid frame approach.

Figure 1. Frame types and relative advantages

Equivalency Requirements

The FSAE competition rules provide detailed requirements

for teams who elect to deviate from the traditional steel space

frame construction [1]. These safety-based requirements are

meant to ensure that the composite monocoque frames are

equivalent to or stronger than the SAE/AISI 1010 steel

frames they are to replace in terms of energy dissipation and

yield and ultimate strengths in bending, buckling and tension.

Figure 2 shows areas on a typical monocoque frame that are

specifically covered by the equivalency rules. Table A1

provides a summary of these rules.

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Figure 2. Monocoque with rear space frame and front

impact attenuator attached.

Laminate Selection and Testing

A honeycomb sandwich structure was selected for the

monocoque construction and the design team relied on

donated materials. These materials included bi-directional

woven carbon/epoxy fabric and unidirectional carbon/epoxy

Cytec CYCOM 5215 Prepreg and Hexcel 5056 Aluminum

honeycomb overexpanded core [8], [9]. This core material

provided superior strength to weight ratio which was

especially important for impact protection [10], [11].

Additionally, FM-300-2 film adhesive and Cytec FM-410-1

adhesive foam were used in the construction [12].

Literature from the Hexcel Corporation listed the properties

of the core material, described common failure modes of

sandwich structures, showed how to analyze panels under a

variety of loading conditions and offered practical

suggestions for sandwich panel construction [13]. Sandwich

panels were conceived, constructed and tested in order to

meet competition rules requirements outlined in Table A1

Two different sandwich configurations were ultimately

chosen based on the testing summarized below. The final

configurations were [F/03/F/core/F/03/F] for the front

bulkhead support region and [F/07/F/core/F/03/F] for the side

impact zones. In these layup descriptions, the ‘F’ refers to the

bi-directional fabric and the ‘0’ refers to the unidirectional

tape and the subscript numeral refers to the number of layers.

The fabric layer preceding the core was a fabrication

requirement that will be presented later. The core was 15.8

mm thick aluminum honeycomb with film adhesive on both

sides.

Flexure Testing

To test the stiffness and strength requirements, two 0.2 m ×

0.5 m flat test panels, [F/03/F/core/F/03/F] and [F/07/F/

core/F/03/F], were constructed for a three point bend test as

shown in Figure 3. Table 1 provides a summary of the test

data. Both configurations were determined to be stiffer than a

single steel tube, and stronger than two tubes as required by

the equivalency rules. The primary failure mode was

buckling of the upper skin due to compression.

Figure 3. Three point bend test of 0.2m × 0.5m flat

sandwich panel

Table 1. Three point bend test results

Perimeter Shear Testing

Next, the two sandwich configurations were evaluated for

perimeter shear strength using a 25.4 mm diameter section.

To accomplish this test, a 25.4 mm steel rod was pushed

through the test samples as shown in Figure 4. As show in

Figure 5, both configurations passed the equivalency

requirements for side impact zone strength.

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Figure 4. Perimeter shear test

Figure 5. Perimeter shear test results

Monocoque Side Impact Validation

Finally, the [F/07/F/core/03/F] configuration was evaluated

against the side impact zone stiffness requirements listed in

Table A1. The vehicle cockpit sidewalls were designed to be

0.48 m tall and the side impact zone (by definition in the

rules) corresponds to a 0.33 m tall section of the sidewall.

Table 2 shows a comparison of EIfor the sandwich panel and

the baseline steel tubing. The first row of data comes directly

from the three point bend test data. The final two rows of data

are derived from the linear relationship between EIand

number of tubes or panel width. From these data, it is clear

that the [F/07/F/core/03/F] configuration meets the EI

requirement for the 0.33 m impact zone and the total side

panel design height of 0.48 m.

Based on the results of both the three point bend and

perimeter shear testing, the final sandwich configuration

chosen for side impact protection was [F/07/F/core/03/F].

Table 2. EI comparison for side impact

Front Bulkhead and Front Bulkhead Support

Details of the design, fabrication and testing of the anti-

intrusion bulkhead and supporting structure were presented in

a previous paper on collision systems for this vehicle [7]. The

final bulkhead design consisted of a 4 mm aluminum anti-

intrusion plate mounted to a “picture frame” consisting of

[F/018/F/core/F/O18/F] members that were 25 mm wide.

Figure 6 shows the bulkhead frame during the construction

phase. Note that the frame members are thin to permit

maximum access to the front of the vehicle for servicing

vehicles components such as the brake and throttle pedal

assembly. Figure 7 shows flexure test data for the front

bulkhead members. The test panel had an ultimate load of

10.5 kN and an EIof 3600 N-m2far exceeding the baseline

tubing values. Despite the obvious overdesign, no further

redesign and retest were attempted due to time constraints.

With the bulkhead member design complete, the next step

was to qualify the mounting points for the impact attenuator

attached to the front of the vehicle. By the rules, each 8 mm

Grade 8.8 bolt had to fail before the attachment to the

bulkhead itself failed. This equated to an out-of-plane tensile

force of 23.0 kN. To meet this requirement, 38 mm diameter

aluminum disks were embedded in place of the core material

at each of the bulkhead corners. Each disk was drilled, tapped

and fitted with a helicoil to receive the impact attenuator bolts

and bonded to both the upper and lower skins as shown in

Figure 6. Use of the aluminum disks enabled both the upper

and lower skins to support the tensile load imposed on the

attachment bolts. During testing the bolt actually suffered a

tensile failure at 30kN before the aluminum disks tore out of

the sandwich structure.

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Figure 6. Front bulkhead under construction with

aluminum inserts for bolts

Figure 7. Bulkhead flexure test data

Table 3. Bulkhead support EI comparison

Next the [F/O3/F/core/F/O3/F] honeycomb sandwich

configuration was evaluated for suitability for the front

bulkhead support region. Using data from the previous

flexural testing shown in Table 1, EI values for the front

bulkhead region were calculated based on a linear

relationship. Table 3 shows a comparison of calculated EI

values that satisfy equivalency requirements.

Monocoque Attachments to Steel

Components

As specified by Formula SAE rules, the main hoop and rear

frame structure must be mechanically attached to the

monocoque with two 8 mm Metric Grade 8.8 bolts bolts at 4

points, for a total of 8 bolts. Also each point must withstand a

30 kN load in any direction. In order to determine the worst-

case loading for the roll hoop attachments, two load cases

were considered: out-of-plane perimeter shear loading and in-

plane pin shear loading. These tests and their results are

described in detail in a previous work [6].

For added strength, inner backing plates of metal were

bonded to the monocoque. The backing plates were made of

3 mm thick steel plate with a bonding area of 14.8 cm2. This

added approximately 15 kN of additional in-plane load

capacity (West System 105 resin thickened with colloidal

silica has been tested to have a shear strength of

approximately 10 MPa [15]), and provided some degree of

“graceful degradation” in place of a rapid, catastrophic

failure.

In order to comply with the rules, each attachment point

between the main roll hoop and the monocoque consisted of a

28-ply monolithic, quasi-isotropic laminate of woven carbon/

epoxy fabric supported on the inner side by a bonded, 3 mm

thick steel backing plate. At each attachment point, this

resulted in out-of-plane strength and in-plane strength in

excess of the 30 kN requirement.

Safety Harness Mounting Points

The proposed safety harness design included lap belts

attached to eye bolts mounted to the monocoque frame that

were required to meet equivalency requirements of 13.0 kN

for the lap restraints and 6.5 kN for the anti-submarine belts.

As detailed in the collision systems paper [7], the first

attempt involved screwing the eye bolts into steel inserts

embedded in a representative honeycomb sandwich [F/O3/F/

Core/F/O7/F] as shown in Figure 8. However this

configuration did not meet the 13.0 kN requirement.

Therefore, a “through bolt” configuration was tested. For this

test, a 3 mm thick circular backing plate 38 mm in diameter

was placed on the outside skin of the honeycomb structure.

The harness mount eye bolts passed through the entire

structure and the backing plate and was held in place by a nut

as shown in Figure 9.

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Figure 8. Schematic of harness mount concept

Figure 9. Through-bolt configuration for harness mount

Figure 10. Harness mount tensile results

The through bolt method proved capable of meeting the test

requirement and experienced a gradual and predictable failure

mode. The back plate slowly deformed as the eye bolt pulled

it and the surrounding skin into the core material. The outer

skin [F/O7/F] never sheared before the test was terminated.

The only failure was crushing of the core material. Based on

these results, the through bolt technique was chosen for the

lap belts and the steel inserts were used for the anti-

submarine belts.

CONSTRUCTION TECHNIQUES

Mold Construction

A female fiberglass mold was fabricated for the monocoque

cockpit layup. First a high density (288 kg/m3) foam male

plug was milled using a Solidworks™ drawing of the cockpit

similar to the one shown in Figure 2. Then the milled plug

was coated with Duratec™ and wet sanded to a smooth

finish. Next a wooden divider panel was made to match the

upper and lower centerline of the plug. Using this divider to

create a flange area, two female fiberglass mold halves were

laid up and cured. Figure 11 shows a photo of the finished

plug with the wooden divider and end pieces attached. The

final mold was gelcoated and had a wide flange with several

holes for easy assembly and disassembly.

Figure 11. Finished plug with wooden dividers ready for

mold fabrication

Prepreg Layup

The following paragraphs detail the construction process used

for the 2012 frame. These procedures include lessons learned

from the 2011 car.

The layup process was completed in three separate steps;

outer skin layup, core insertion and finally inner skin layup.

Each of these three steps are described in the following

paragraphs.

To begin the layup process the mold halves were bolted

together and treated with the release agent Frekote™. An

external wooden support frame was bolted to the mold to

prevent warping and provide stability during the layup. Four

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large paper patterns were cut to represent the floor, sides, side

impact region and top of the cockpit. These patterns were

used to aid in cutting the carbon fiber prepreg material

required for the outer skin.

First one fabric layer and three unidirectional layers were laid

into the mold. All pieces were hand laid into the mold taking

special care to work out air bubbles and prevent “bridging”

across all features. In the side impact region, four additional

unidirectional layers were applied. Lastly a fabric layer was

used to cover all of the unidirectional tape, thus providing

biaxial stiffness on both sides of the composite skin. The

majority of the outer skin consisted of [F/O3/F] and the side

impact regions were [F/O7/F]. Finally ten fabric “patches”

were applied to reinforce the monocoque attachment points as

required by the test results.

The importance of the final fabric stiffening layer cannot be

overstated. In an earlier layup this fabric layer was omitted

[F/O3] and the results were catastrophic. Upon removal of the

outer skin from the mold for inspection, the skin buckled and

cracked along the direction of the unidirectional fibers in

several areas. This damage occurred even with the most

careful handling and rendered the outer skin unusable. In the

next attempt, the final fabric layer was used and this

prevented skin buckling. It is recognized that the need this

fabric layer could be avoided if the outer skin and core were

laid up together. However, for novice fabricators, it is critical

to inspect the layup between each process. Additionally, the

extra fabric layer enhances the torsional rigidity of the

finished monocoque.

(a). passing the vacuum bag through the monocoque

(b). sealing the vacuum bag back on itself

Figure 12. Cross-section schematic illustrating double

bagging technique. For clarity, only the monocoque and

vacuum bag are pictured.

Figure 13 shows a picture of the sealed layup going into the

autoclave. The curing process included an 8 hour debulking

period at 28 inHg and room temperature to remove voids

from the layup followed by a 2 hour cure at 127°C [9]. Note

that the autoclave was only used for heat and vacuum as the

part was not cured under applied pressure.

Figure 13. Vacuum bagged layup going into the

autoclave

The next step involved laying in the core material. The paper

templates were used to rough cut the film adhesive and core

and these pieces were laid into the cured outer skin. The film

adhesive was cut away from the reinforced monocoque

attachment points to enable easy removal and trimming of the

core material. This was necessary as these points will become

the 28-layer monolithic regions as required by the previous

testing. Foaming adhesive was used to join the core sections

and the layup was double vacuum bagged. This time the

vacuum was set to 15 inHg to avoid crushing the core

material. Cure time was 2 hours at 127°C. Upon removal

from the autoclave, the core was trimmed in preparation for

the inner skin layup.

Prior to laying in the inner skin, aluminum “hard points” with

helicoil inserts were pounded into the core for attaching

critical components such as the brake pedal mount, shock

mounts and steering column mounts. This was a required

improvement over the 2011 design in which these critical

components were only bonded to the finished monocoque

rather than being bolted in place. One of the required safety

checks at the competition includes braking hard enough to

lock all four wheels from a moderate speed. During this test,

the brake pedal mount broke away from the monocoque floor

due to a failure of the secondary bond (in the 2011 car).

While the car safely came to a stop, extensive redesign was

required to mechanically attach a new brake pedal mount to

the floor. This precluded the team from competing in three of

the four driving events. In 2012 all attachment points were

bonded and bolted for added strength and ease of repair.

The inner skin layup was the last step in the monocoque

fabrication process. Reinforcement patches were placed over

roll hoop and suspension attachment points to complete the

monolithic structure. Then the core was covered by a layer of

film adhesive, bi-directional fabric, three layers of

unidirectional tape and a final layer of bi-directional fabric.