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LIBRAR‘s’ i
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. University (

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This is to certify that the

thesis entitled

KINETIC MODELING AND SEAT FACTORS RELATING TO
MANIKlN-SEAT INTERACTION IN AUTOMOBILES

presented by

Nathan J. Radcliffe

has been accepted towards fulfillment
of the requirements for

MS Engineering Mechanics

 

 

degree in

Major professor

8-23-01
Date

 

0-7639 MS U is an Affirmative Action/Equal Opportunity Institution

 

PLACE IN RETURN Box to remove this checkout from your record.
TO AVOID FINES return on or before date due.
MAY BE RECALLED with earlier due date if requested.

 

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8/01 cJClFlC/DatoDuepes-sz

 

 

KINETIC MODELING AND SEAT FACTORS RELATING TO
MANIKIN—SEAT INTERACTION IN AUTOMOBILES

By

Nathan J. Radcliffe

AN ABSTRACT OF A THESIS

Submitted to
Michigan State University
in partial fulfillment of the requirements
for the degree of

MASTER OF SCIENCE

Department of Materials Science and Mechanics

2001

Dr. Robert Hubbard

ABSTRACT

KINETIC MODELING AND SEAT FACTORS RELATING TO
MANIKIN-SEAT INTERACTION IN AUTOMOBILES

By

Nathan J. Radcliffe

Automotive seats are integral components in the interactions between occupants
and vehicles, and ultimately the road surface traveled. Seats are important in determining
the comfort of the ride, and therefore the user’s impression of the total vehicle. The
occupant’s static deflected position, supported posture and comfort may be related to a
set of quantitative values known as Seat Factors. Seat Factors are quantitative measures
and calculated values of structural properties for complete seats related to biomechanical
interactions. Using specific procedures, measurable Seat Factors can be attained which
relate to the physical properties of each seat pan. These measured quantities may then be
used in a kinetic model, which calculates additional Seat Factors that can be used to
predict the static deflection and rotation response of the buttocks and thigh region in
contact with the seat pan, due to imposed loads at the hip joint center and knee. This
thesis provides information regarding the procedures used to determine Seat Factors,
assumptions used in modeling, kinetic models and components, model validation

techniques, applications and future work.

Copyright by
Nathan J. Radcliffe
© 2001

ACKNOWLEDGEMENTS

I would like to thank Dr. Hubbard for positively influencing my engineering
career over the past six years. Your positive guidance, mentoring, fi’iendship, and
support have guided me in the right direction through the often tough but always

encouraging years of engineering school at Michigan State University.

Thank you to my master’s committee members, Dr. Reid-Bush, and Dr.
Rosenberg. Your help and guidance has helped this project and my education progress

successfully through both my undergraduate and graduate programs.

I would like to thank my family for pushing me to succeed from the very
beginning. Mom, Dad, Isa, and Eliott, I thank you for all of your loving support. And
finally, I would especially like to thank my fiancé Terri for her loving patience and
support without which I would be further challenged to succeed in this project and

beyond.

TABLE OF CONTENTS

LIST OF TABLES VII
LIST OF FIGURES Ix
NOMENCLATURE XI
INTRODUCTION 1
EXPERIMENTAL DATA ACOUIsrrION METHODS 2

SAE 1826 W’ Seating Manikin 2

ASPECT Seating Manikin 5
KINETIC MODELS a
Model I. Two-spring kinetic model with support forces under the H-pt. and knee. 6

Model I]. Two-spring kinetic model with single spring constant, support forces under the H-pt. and
thigh at a fixed distance ETSL from the H-pt. 7

Static Mathematical Analysis for Physical Simulation of Model 11 using a Dynamic Solver Software 8
Results from Model I] 11

Model III. Two-spring kinetic model with variable spring constant support forces under the H-pt.

and thigh at a variable distance ETSL from the H-pt. 13
Results from Model III 14
Model IV. Coupled force and moment kinetic model 16

Static Mathematical Analysis for Mathematical Simulation of Model IV using MS Excel Solver to
Optimize Calculated Seat Factors 17

Mathematical Modeling Methods 19
Step 1. Experimental Data Input 19

Step 2. Static Reaction Force and Moment Calculation 20
Step 3. Optirnizatim of Modified Taylor Series-Based Seat Factor Matrices 21
Step 4. Solving Equations to Determine Simulation Results Based on Calculated Seat Factor Matrices

23

Step 5. Graphical Analysis of Optimization and Simulatim Results 24
Step 6. Occupant Positim Approxrmat' ions 29
Results and Conclusions from Model IV 31
CONCLUSION 39
Future Work 40
APPENDICES
APPENDIX A 45
APPENDIX B 168
APPENDIX C 170

REFERENCES 183

vi

LIST OF TABLES

Table 1. Input data acquired fiom experimental incremental loading of the rmnikin is
shown in white font within dark cells, along with calculated values based on
experimental data shown in grey. 20

Table 2. Individual Deflection-Rotation Solver for Occupant Position Approximations. 29

Table 3. Average H-pt. deflection, knee deflection and relative thigh rotation difference
across experimental loading range for linear, second, and third order Seat Factor

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

Matrices. 32
Table 4. RMS and Percent Error vs. Order of Taylor Series Expansion. 37
Table L-l . Seat Key for seats analyzed within the Seat Factor Solver Spreadsheet.____47
Table A-1. Seat A SFS Analysis Worksheet. 48
Table A-2. Seat B SFS Analysis Worksheet. 52
Table A-3. Seat C SF S Analysis Worksheet. 56
Table A-4. Seat D SFS Analysis Worksheet. 60
Table A-5. Seat E SFS Analysis Worksheet. 64
Table A-6. Seat F SF S Amlysis Worksheet. 68
Table A-7. Seat G SFS Analysis Worksheet. 72
Table A-8. Seat H SF S Analysis Worksheet. 76
Table A-9. Seat I SFS Analysis Worksheet. 80
Table A-lO. Seat J SFS Analysis Worksheet. 84
Table A-1 1. Seat K SFS Analysis Worksheet. 88
Table A-12. Seat L SFS Analysis Worksheet. 92
Table A-13. Seat M SFS Analysis Worksheet. 96
Table A-14. Seat N SFS Analysis Worksheet. 100
Table A-15. Seat 0 SFS Analysis Worksheet. 104
Table A-l6. Seat P SF 8 Analysis Worksheet. 108

 

vii

Table A-17

Table A-1 8.
Table A-1 9.
Table A-20.
Table A—21.
Table A-22.
Table A-23.
Table A-24.
Table A-25.
Table A-26.
Table A-27.
Table A-28.
Table A-29.
Table A-30.

. Seat Q SF S Analysis Worksheet.

 

Seat R SFS Analysis Worksheet.

 

 

Seat S SFS Analysis Worksheet.
Seat 4 SFS Analysis Worksheet.

 

Seat 8 SF S Analysis Worksheet.

 

Seat 9 SF S Analysis Worksheet.

 

Seat Neon/ASPECT SFS Analysis Worksheet.

Seat TandC/ASPECT SFS Analysis Worksheet.

 

Seat Saturn/ASPECT SFS Analysis Worksheet.

 

 

Seat TandC/J826 SF S Analysis Worksheet.

Seat LHtan/ASPECT SF S Analysis Worksheet.

 

 

Seat Saturn/1826 SFS Analysis Worksheet.

 

Seat Neon/J826 SFS Analysis Worksheet.

 

Seat LHtan/J826 SFS Analysis Worksheet.

 

viii

112
116
120
124
128
132
136
140
144
148
152
156
160
164

LIST OF FIGURES

 

 

Figure 1. SAE 1826/0scar manikin (Unloaded - Loading Step 1). 4
Figure 2. ASPECT manikin (Unloaded). 5
Figure 3. Two-spring kinetic model with support forces under the H-pt. and knee. 6

Figure 4. Two-spring kinetic model with single spring constant, support forces under the H-
pt and thigh at distance ETSL fi'om the H-pt. 7

Figure 5. Descriptive two-dimensional figure of H-pt.-knee rotation about the ETSL for
similar triangle-based analysis of the ETSL. 9

Figure 6. Descriptive two-dimensional plot of H-pt.-knee spring constant linear magnitude
relationship for calculation of ETSS at the ETSL. ' 10

Figure 7 a-d. Examples of Working Model results fi'om physical simulations of Model H.
Plots indicate good experimental-simulation results for a linear seat (Seat G) in

Figures 7a and 7b and poor experimental—simulation results fi'om a non-linear seat
(Seat A) in Figures 7c and 7d. 11

Figure 8. Two-spring kinetic model with variable spring constant support forces under the
H-pt. and thigh at a variable distance ETSL from the H-pt. 13

Figure 9 a-d. Examples of Working Model generated results fi'om physical simulations of
Model III with results of Model 11 shown for comparison. Plots indicate
experimental-simulation results for a linear seat (Seat G) in Figures 9a and 9b and
experimental-simulation results from a non-linear seat (Seat A) in Figures 9c and
9d. ' 15

Figure 10. Coupled force and moment kinetic model. 16

Figure 11 a,b. Plots for comparison of experimental results (diamonds) to simulation results
based upon third order Taylor series expansion matrices (line) (Seat A). 25

 

Figure 12. Thigh position as a fimction of incremental H-pt. and knee loading. The
graphical point of thigh rotation (zero deflection) during H-pt. loading is
approximately 225mm fi'om the H-pt. along the thigh for this seat (Seat A). 26

Figure 13 a,b. Third order surface models of reaction force and moment as a function of H-
pt. deflection and relative thigh rotation (Seat A). 28

Figure 14. Interpolated and extrapolated H-pt. force surface Plot, shown with extrapolation
and interpolation beyond the path of experimental data (Seat 0). 33

Figure 15. Two-dimensional contour of the SAE 1826 Manikin [4]. A ‘/4 inch, 2-D plastic
template with rods extending from the H-pt. was used as a reference for data
acquisition by Hubbard and Gedratis [1] and Hubbard and Radclifl‘e [6] within
this project [4]. 35

Figure 16. Proposed two-dimensional linked system of reaction forces and moments (dark
arrows) with input loading (white arrows) at crucial joints 42

Figure C-l . Descriptive two—dimensional figure of H-pt.-knee deflection and rotation for
analysis of the ETSS. 182

NOMENCLATURE

Arabic Symbok and Acronyms
ETSL = efl‘ective thigh support location
ETSS = effective thigh support stimress
F = generic reaction force under the H-pt (could be RH or F")
Fn = simulation reaction force under the H-pt.
FH = input force at the H-pt.
FK = input force at the knee
FT = input force at the thigh
n K fl, = nm order reaction force-deflection stifiness coeficient
n K ,9 = nth order reaction force-rotation stimtess coeficient
n K m 6 = 11‘” order reaction moment-deflection stiffness coeflicient
" K m 9 = 11‘11 order reaction moment-rotation stifliress coefficient
Kn = spring constant of the seat under the H-pt.
KK = spring constant of the seat under the knee (physically non-existent)
LK = distance from the H-pt. to the knee along the thigh

xi

Mn =

SF =

Greek Symbob

5" =

.9" =

9 THIGH =

large male (90th percentile for height and weight, 180 lbs.)
generic reaction moment around the H-pt (could be Mu or Mn).
simulation reaction moment around the H-pt.
experimental reaction moment around the H-pt.

mid male (50* percentile for height and weight, 170 lbs.)

experimental reaction force under the H-pt.

small female (10til percentile for height and weight, 115 lbs.)

relative deflection of the H-pt. (same as AH)

relative rotation of the thigh (referenced to the 2-D template)

relative deflection of the H-pt. (referenced to 2-D template)
relative deflection of the knee (referenced to 2-D template)

thigh angle referenced to horizontal

xii

INTRODUCTION

The pln'pose of this investigation was to develop an experimental data
interpretation technique coupled with a kinetic modeling method to quantitatively
describe and evaluate the mechanical properties of automotive seats. These quantitative
measures and calculations are known as Seat Factors [lfit was theorized that ifa
modeling method could be developed which produced similarresults to data acquired
experimentally when identical loading procedures were followed, the model and a
calculated set of Seat Factors were validated and could be used to define the mechanical
characteristics of a particular seatét‘The evolution of this project included several different
kinetic models, and their validatioh using experimental results fiom 30 production and
prototype automotive seats, which is discussed in upcoming sections. For the initial
kinetic modeling, static equations were derived to solve for calculated Seat Factors, and
physical modeling was attempted for the purpose of validation using Working ModelTM
(WM) [2]. {I Working Model“ is a two-dimensional, dynamic simulation software, which
has the abrlrtjito predict motion of rigid bodies with imposed forces and constraints.
Though this physical seat factor validation technique provided much insight, the most
successful kinetic modeling method involved experimental data acquisition coupled with
mathematical analysis using Microsoft (MS) Excel Solver [3] to optimize modified
Taylor series expanded, two by two descriptive seat factor matrices based upon the

experimental response of the manikin loaded incrementally into the seat]

Traditionally, seating design is based upon analysis of comfort, occupant posture,

contact surface pressure distribution, loaded manikin behavior, and software model

positioning timing the computer-aided design process. Though these techniques provide
designers with valuable insights into the interactions between people and seats, these
methods do not lead to a complete, quantitative set of Seat Factors that defines the
structural characteristics ofthe seat}; was theorized that a simplified, condensed set of
Seat Factors coupled with a kinetic modeling approach could be used as a predictive tool
for occupant positioning, and may also be used to correlate quantitative Seat Factors to
subjective responses in terms of comfort. 6he latter of the two proposed applications will

_ ”may (‘
be considered for future work, and will not be investigated within this thesis. ‘1‘ 1

Previous work in seating biomechanics and kinetic modeling by team members of
the Biomchanical Design Research Laboratory at Michigan State University has lead to
the definition of Seat Factor data acquisition techniques [1] and the development of the
ASPECT manikin [5] through support of the ASPECT program. The ASPECT Manikin
is a replacement for SAE’s J826 "Oscar" maniklnEThis manikin has an articulated back
section that can be used to take various manikin-seat interaction measurements including
hip joint center location, thorax, pelvis and thigh angle, seat back and seat pan angle, and

lumbar prominencei'i

EXPERIMENTAL DATA ACQUISITION METHODS

SAE J826 “Oscar” Seating Manikin

Gil a previous work by Hubbard and Gedraitis [l ], an experimental technique for

measured seat factor data acquisition with the SAE J 826 [4] manikin was used. This

technique is presented in detail within SAE Paper Number 1999-01-09'5’31‘ but can be

._..——J

Ar"

summarized in the following steps:

1.

Using the SAE J826 “Oscar” seating manikin (shown in Figure 1), incremental loads

were measured at the H-point (H-pt.) and knee axes.

Note: H-point (H-pt.) is a term commonly used within the automotive industry, which
relates to the location of the H-pt. axis on seating manikins when a specific loading
procedure is followed [4]. The H-pt. axis is analogous to the hip joint center (HJC) of
a human. References to the H-pt. within this report refer to the location of the
physical H-pt. axis on the mnikin, and not the position of the H-pt. axis according to

the established H-pt. procedure [4].

This was accomplished by suspending the seating manikin along the H-pt and knee
axes by tensile lines with forces measured using voltage-output load cells.
Incrementally loading the H-pt axis while holding a constant, minimal knee load, and
subsequently, incrementally loading the knee axis with no additional applied H-pt
load investigated a range of loads similar to a person sitting down into a seat. The

measured loads could then be known for that particular incremental loading sequence.

  

 

 

I” u“. ,doJ *'.-_

Figure l. SAE J826/Oscar manikin (Unleaded -
Loading Step 1)

f”.

I, ; (o;

II(A reference position in each seat was measured using the 2-D plastic SAE J826
Iemplate [1,4])The manikin without additional weight was then placed on top of the
seat pan surface with the rear of the manikin seated firmly against the lower portion
of the seatbackCPrior to measuring the position of the loaded SAE 1826 manikin in
the seat, the torso/back section of the manikin was adjusted to vertical for each , wa
incremental load. This minimized interaction between the back of the manikin and
the seat back, and ensured consistency between trials. Known loads were
incrementally added, and the positions of the H-pt and knee axes were measured afier
each applied load. H-pt. axis measurements were made on the left and right side by
measuring the position at the tip of a rod extending out fiom the H-pt. axis center.
Left and right H-pt and knee axis measurements were measured at the same distance
from the vertical plane of symmetry of the manikin. By averaging the left and right
measures, the variation in level on each side of the manikin was averaged. Average

measures, which corresponded to deflections at the plane of symmetry of the manikin,

were used for simulation data input purposes. Measurements of H-pt and knee
positions could be acquired using a motion analysis system or any method of vertical

position measurement with a fixed reference i.e. floor [1].

ASPECT Seating Manikin

Data acquisition using the prototype ASPECT manikin [5], butt-thigh segment
(ABT), was accomplished in a similar manner as the procedure discussed for the SAE
1826 manikin. Only the butt-thigh region of the ASPECT manikin was used for data
acquisition? order to remove interaction between the back section of the manikin and , , ‘
the scat [6T Data acquisition techniques and results ofASPECT manikin testing are
presented in detail within Appendix B: Seat Pan Load-Deflection Measurements For
Determination of Seat Factors with the ASPECT Butt-Thigh (AB7) Segment of the
ASPECT Manikin Prototype at MSU and the SAE J826 Manikin by Robert Hubbard and
Nathan Radcliffe. Results of ASPECT manikin data acquisition and simulation are

discussed with comparisons to the SAE 1826 within Appendix B.

 

ABT frOI'lt _ f 3 w, p: .> _' d

 

 

Figure 2. ASPECT manikin (Unleaded)

KINETIC MODELS

During the course of work on this project, four separate kinetic models were
successively developed. Each of the models are discussed in this section with varying

levels of description based upon their relevance to the final outcome of the project.

Model I. Two-spring kinetic model with support forces under the H-pt.
and knee.

Slider prohibits horizontal translation of the H-pt.

Flt

 

H-pt

 

 

Figure 3. Two-spring kinetic model with support forces under the H-pt. and knee

Model I, a two-spring kinetic model with support forces under the H-pt. and knee,
was the first model developed. It separated the complex, distributed load interaction
between an automotive seat and a manikin into two input forces and two support forces at
the H-point and knee. This model was a simple way to represent experimental behavior,
as it was trivial to calculate the spring constants under the H-pt. and knee with a
measured force-deflection response. It was realized prior to the onset of analysis and
modeling that a support force directly under the knee did not allow the thigh to rotate

about a point between the H-pt. and knee as observed in measurement by Hubbard and

Gedraitis [1]. In order to allow for realistic rotation behavior of the thigh in a kinetic
model, the support must be positioned at some distance known as the effective thigh
support location (ETSL) away fi'om the H-pt. along the thigh, which lead to the

development of Model II.

Model II. Two-spring kinetic model with single spring constant,
support forces under the H-pt. and thigh at a fixed distance ETSL from
the H-pt.

 

 

1 Slider prohibits horizontal translation of the H-pt.F
K
FH
L
K . | Knee
H-Pt

’ 0; ETSL ETSS

;_ Kn
—

Figure 4. Two-spring kinetic model with single spring constant, support forces under the H-pt
and thigh at distance ETSL from the H—pt.

In Model H, the first support force was located at the same point under the H-pt.
as seen in Model I. This model differed fiom Model I, in that the second support force
was located at some distance known as the eflective thigh support location (ETSL) away
fi'om the H-pt., along the thigh In designing this kinetic model, it was hypothesized that
the second support would exist at the point of zero deflection (ETSL; thigh pivot point
referenced to original thigh angle) during H-pt. loading: Based upon experimental

measurements [1] and calculations, the distance fi'om the H-pt. to the point of rotation

along the thigh (ETSL), rennined relatively constant during H-pt. loading and typically
stabilized at larger imposed loads. Therefore, a single ETSL was chosen based upon a t
middle point in the incremental H-pt. loading range to simulate the hill-range, 9 '
incre -loaded response of the manikin. The single ETSL used for all imposed
loads in simulation was calculated based upon the point of rotation for loads of494N on

the H-pt., and 29N on the mi?

Based upon experimental data acquired by Hubbard and Gedraitis [1], one spring
constant for each of the two support springs was calculated to represent the behavior of
the seat throughout all imposed loads in the software simulation. These spring constants
were referred to as K" (under H-pt.) and ETSS (under thigh) or effective thigh support
stiflhess. The H-pt. spring constant (Kn) was calculated based upon the relative force-
deflection response at 494N on the H-pt., and 29N on the knee. The ETSS used for

simulation was an average of the calculated ETSS during the full range of incremental
knee loadings.

Static Mathematical Analysis for Physical Simulation of Model H using a Dynamic
Solver Software

Using static analysis, the following equations were derived to calculate individual
Seat Factors relevant to the kinetic simulation of Model I] (and Model 111) below. The
spring constants KH and KK were based upon relative, incremental characteristics of the
force-deflection response of a particular seat. Each of these spring constants were the
incremental input force (in this case referenced to the SAE 2-D template [4], which was
assumed to have negligible mass) divided by the relative deflection referenced to the

initial position of the H-pt. and knee as established by the SAE 2-D template.

The effective thigh support location (ETSL) was at the point of zero deflection
(instantamous pivot) during H-pt. loading with constant nominal knee loads, referenced
to the initial position of the SAE J826 template [4]. When the H-pt. deflected into the
seat and the knee elevated during incremental H-pt. loading, there was always a point of
zero deflection. Calculation of the ETSL was done for H—pt. loading only. The ETSL
used dining H-pt. loading was assumed to be the constant ETSL for incremental knee

loading. This point was calculated by similar triangle-based analysis as derived below.

 

 

— Initial Position
— Position after H-pt. Loading

Figure 5. Descriptive two-dimensional figure of H—pt-knee rotation about the ETSL
for similar triangle-based analysis of the ETSL

LK _AH-AK
ETSL AH

 

AH*LK

ETSL:—
AH—AK

(1)

In order to calculate the efi‘ective thigh support stifi‘ness at a point of support
between the H-pt. and knee, the equation derived below using proportional analysis of

similar triangles was used. This ETSS equation was only valid during knee loading, but

the average ETSS value calculated from experimental knee loading was used as a

constant for incremental, simulated H-pt. loading.

Assumption: Linear change in spring constant magnitude between H-pt. and knee

  
 
  
 

Magnitude of K“

 

H-pt.

 

+
Lx AH- AK

 

ETSL

Figure 6. Descriptive two-dimensional plot of H-pL-knee spring constant linear
magnitude relationship for calculation of ETSS at the ETSL.

K” —Erss _ K” -KK
ETSL LK

 

ETSS: K” — ETSL“? 'K") (2)
K

 

Equation 2 for ETSS was based upon the assumption that the spring constant at
the location of the thigh support was linearly related to the spring constants under the H-
pt. and knee (virtual, as the knee of the manikin extends beyond the support of the seat)

according to the effective location of the thigh support between the H-pt. and knee along

the thigh. Typically, automotive seats decrease in stiffness as you move from the support

under the H-pt. toward the fiont of the seat underthethigh The constant value used for

 

KH was the last calculated spring constant during H-pt. loading with constant nominal
knee loads.

An alternative equation for determining ETSS proposed after Model II analysis,

which was not fully investigated due to the success of Model IV, is given in Appendix C.

10

Results from Model H

Whenasinglespringconstantwasusedforeachspringineth[2]trial,

along with a constant ETSL, the results of all physical simulations for Model H yielded

linearly varying deflections at the H-pt. and knee as a function of applied load. Two

sample sets of typical results for Model H are shown below in Figure 7 a-d. Simulation

results for an approximately linear seat (Seat G) and a stiffening, non-linear seat (Seat A)

are shown. These two seats represent good (Seat G) and poor (Seat A) experimental-

sirnulation correlations and they were indicative of the range of results of Model H.

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

mefi) KnoeDclIectlonIG)
ms 235 335 435 585 see 0 50 100 150 200 250 300
o . A 40
E .5- +ModaIIl v; 20‘ +Modalll
: -1: \ +Exp : o j—\.—I- , +59
E ' E \
é -20 5 -20 \
-25
g ,0 x t g 4° \.
-35 l -80 i
.40 ‘ a0 1
H-pt.Loadino Fema(N) KneeLoad‘ng Force (N)
Figure 7a Figure7b
WWNW KnceDcsectlontAI
135 235 ass 485 585 ass 0 so 100 150 200 250 300
A 0 ' A 30
-5 +M°°°'" ‘ 20. +Modalll
3 _ \4‘\ g 10..
I 10 \ \ +5“) I O \ +531)
E45 ‘Q . £540 \ \ 1
cm \ A? .20 \ \ l
-25 -30 \ \_
so \ \‘
'35 H-ptl-owimFomem) - KnceLoadanoroerN)
Figure 7c Figure 7d

Figure 7 a-d. Examples of Working Model results from physical simulations of Model H. Plots
indicate good experimental-simulation results for a linear seat (Seat G) in Figures 7a and 7b and
poor experimental-simulation results from a non-linear seat (Seat A) in Figures 7c and 7d.

 

For seats that exhibited a linear force-deflection response in experimental testing
(as with Seat G), simulation results were similar to the measured responses. However,
seats that exhibited non-linear (typically stiffening) forceodeflection responses during
experimental manikin loading (as with Seat A) did not yield results that were acceptably
similar between the experimental response and kinetic model physical simulation. This
was due to the choice of a single spring constant calculated fiom the force-deflection
results of a middle loading step (494 N of H-pt. force and 29 N of knee force). A single
spring constant seat factor could not effectively simulate the soft behavior of the
stiffening seats when small, initial loads were applied. In Figure 7c, the large load
stifliless under the H-pt. of Seat A is similar for both the experimental data and physical
simulation, as indicated by approximately parallel lines between 385 N and 585 N OfH-

pt. loading force. The offset difference rs due to the imbility of a single (linear) spring

constant to effectively simulate a stiffening (non-linear) seat with soft properties at small, .

impqsed loads- ..

Further downward deflection of the H-pt. during H-pt. loading in experimental
results caused the knee to deflect upward by means of rotation prior to the onset of knee
loading. This phenomenon accounted for the offset in knee deflection during knee
loading. In other words, differences between Model H simulation and experimental

results of H-pt. deflection influenced Model H physical simulation of knee deflection.

12

Model HI. Two-spring kinetic model with variable spring constant
support forces under the H-pt. and thigh at a variable distance ETSL
from the H-pt.

Slider prohibits horizontal translation of the H-pt.

 

H-pt

 

ETSS(FH)

 

Slider allows for variation of ETSL

Figure 8. Two-spring kinetic model with variable spring constant support forces under the H-pt.
and thigh at a variable distance ETSL from the H-pt.

Analysis of Model H simulations suggested that a variable spring constant model,
with variable ETSL would likely produce better results since Seat Factors varied with
incremental loads. Model HI was investigated extensively with physical simulations in
Working Model”. Using WM’s data import abilities, simulations were run with variable
spring constants KH(FH), ETSS (FH, Fk), and ETSL (FH ) as a function of the applied loads
on the H-pt. and knee. The choice of variable spring constants and a variable ETSL
required data exchange between WM and a spreadsheet, along with the construction of a
sliding thigh support under the thigh Each of these factored dramatically into the

complexity of this physical model.

Mathematically, the same equations presented in the description of Model H were
applied directly to the physical simulation of Model IH. The difference existed in the

selection of variable K", ETSL and ETSS as fimctions of incremental loads. Since KH

,.-

13

l “ F“

and ETSL were calculated based only upon experimental deflection results of H-pt.

-m-..’_—_~_.r-. .~_

loading with constant knee loads, the final coefficient of each of these terms was used as
a constant coefficient during knee loading. Comparably, for the ETSS, the first -
coefficient calculated fi‘om incremental knee loading with constant H-pt. loading was i

used to represent the ETSS durinng-pt. loading with. constant knee loading.

fl’-‘—-—‘——-v~- —_ I ""‘

Results from Model III

When variable spring constants were used for each spring in each trial, along with
a variable ETSL, the results of all physical simulations for Model 1H yielded non-linear
deflections at the H-pt. as a fimction of applied load. Two sample sets of typical results
for Model ID is shown below in Figure 9 a-d. Simulation results for an approximately
linear seat (Seat K) and a stiffening, non-linear seat (Seat A) are shown. The Model HI
response under the H-pt. was acceptably similar to experimental deflections for each of
the seats. Problems occurred in simulating the deflection of the knee or rotation of the

thigh.

l4

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

H-potrlDeIledimoo WWMW
135 235 335 435 535 335 0 50 100 150 200 250 300
A 0 T +Modalll A 2
g -5 +Exp g 104%
I‘m V 04 - \ .
E +Modalm g \ \o
g -15 g -10 \ J
'2° 3: +Modclll \ l
.25 40‘»— +Em x
40 .50.L_
H-DLLoadimForcam +Modellll KmaLoadingFomem)
Figure9a
H-poIruDeIIectIonw
185 285 385 485 585 685
g'K ' t ' +W"
E -10 X\ +Exp S 4
\
v 15 A +ModaIlll L. ‘
g .20 \I E - .
5 a s -20
-30 w— 8 .30.
-35 g 40«
~40 4] .50.L_
WWWFUOBM +m'" KneeLoad'noForce(N)
Figure9c Figure9d

Figure 9 a-d. Examples of Working Model generated results from physical simulations of Model
111 with results of Model 11 shown for comparison. Plots indicate experimental-simulation
results for a linear seat (Seat G) in Figures 9a and 9b and experimental-simulation results from a
non-linear seat (Seat A) in Figures 9c and 9d.

Though Model IH could effectively represent H-pt. deflection during physical
simulation, significant differences occurred between simulation results and experimental
data for knee deflection. The differences were attributed to the dual-dependency of thigh
rotation and knee deflection upon the ETSL and ETSS. If either of these parameters
contained errors, the response was typically unacceptable. Since the equations presented
for ETSL and ETSS could not be used directly for all types of loading as previously

discussed, error due to assumptions of constant ETSS for H-pt. loading, and constant

15

ETSL for knee loading caused the rotational response of the thigh to be less than

acceptable.

Model IV. Coupled force and moment kinetic model

Slider prohibits horizontal translation of die H-pt.F

 

 

H-pt

 

Figure 10. Coupled force and moment kinetic model

Model IV was a simplified development from Models II and HI for modeling the
support of an automotive seat. Mode] IV used only two Seat Factor terms to describe the
structural properties of each seat. This model represented the support structure of the seat
using a reaction force-under the H-pt. and a coupled moment about the H-pt. axis of the
seating ' The generic support force Fjfilied on the same basic assumptions as
spring constant-based support force (Kn) in the previous models (I-III). Model IV
differed from Models I through 111 since the behavior of the seat between the H-pt. and
the knee was described by a reaction moment M, and a rotation of the thigh between the
H-pt. and knee. The generic reaction moment about the H-pt. axis M, was coupled to the
support force F. This approach exchanged the two previous Seat Factors, ETSL and

ETSS for a Seat Factor M based on the moment angle response. By reducing the number

16

of calculated Seat Factors, the model and corresponding Seat Factor calculation
procedures were simplified. The coupled force and moment system of Model IV was
extremely effective for modeling the effects of the seat on manikin posture. However,
the separation ofthe moment into a force and distance increased the level of

understanding of thigh interaction with a seat.

Initially, Model IV was investigated using a similar physical kinetic modeling
approach within WM as Models I through III. During initial attempts to simulate the
response of the seat using Model IV, significant dynamic instabilities were present within
the static physical modeling in WM. [It was then realized that the use of a dynamic solver
to determine static results was not an ideal approach to solving the proposed problem.

An important point to be made is that this is currently a static investigation into the force-

deflection characteristics of automotive seats. A dynamic investigation could be pursued N, .fl

in future work based upon a successful static analysis of mechanical seat characteristics.

Due to these dynamic instabilities, a mathematical modeling approach using static
analysis techniques was investigated. As will be shown from the results of Model IV, the
mathematical modeling approach provided the best results during simulations.
Mathematical modeling is also software independent, which increases applicability. For
these reasons, this mathematical approach will be the basis of material discussed in the

remainder of this thesis.

Static Mathematical Analysis for Mathematical Simulation of Model IV using MS
Excel Solver to Optimize Calculated Seat Factors

Model IV represents two input forces at the H-pt. and knee, and two reactions: a

coupled reaction force at the H-pt. axis and reaction moment around the H-pt. axis. To

17

efl‘ectively simulate the response of an automotive seat to two variable input forces, the
reaction force and moment were calculated using the force-deflection results and physical
parameters of the experimental data acquisition process. Calculation of Seat Factors was
accomplished using the MS Excel Solver optimization tool and a MS Excel [3]
spreadsheet designed for this analysis that will be referred to as the Seat Factor Solver

(SFS).

The processes for determining calculated Seat Factors related to Model IV, along
with mathematical simulations using the calculated seat factor parameters, are described f),
in the following sections( This procedure used modified Taylor series expansion maize?)
to approximate the behavior ‘of the seat during experimental data acquisition. Modified i ‘
Taylor series equations were used to successively decrease the difieremes between I
experimental data and simulation results with the addition of error correcting, higher

order terms. These Taylor series expansions are modified in that the 0‘” order,

linearization term is neglected, along with terms which are coefficients multiplied by

both variable terms (K*[6 " * .9" ])(see equations 5 and 6). First, second, and third order

modified Taylor series expansions were investigated within the Seat Factor Solver to
{r/educe the RMS difl‘erenceaéween the experimental and simulation results. The
ms of each modified Taylor series term increased the complexity of the kinetic
model, but potentially reduced the differences between the experimental and model
responses. Following the computation of the modified Taylor series seat factor matrices,
several tables and plots were generated to graphically interpret the matrix-based

simulated response of the kinetic model to imposed loads. Each of these steps are

documented within this section.

18

Mathematical Modeling Methods

The objective of this mathematical modeling procedure is to determine a set of
two simulation equations (F .. and MD) with optimized parameters

("K13 ,nKfi, ’nKnuS ”KM ), which represent the behaviors ofautomotive seats based

upon experimental results and can be used to simulate these behaviors.

3

Step 1. Experimental Data Input

Prior to optimization and simulation, experimental data must be acquired
according to the experimental data acquisition procedure documented by Hubbard and
Gedraitis [l] and Hubbard and Radcliffe [6]. Data acquired previously by Hubbard and
Gedraitis [1] was'used for simulation with Models 11 through IV. An additional set of
data acquired by Hubbard and Radcliffe [6] (Appendix B) was added to Model IV
mathematical simulations using the SFS to investigate the response of the ASPECT vs.

SAE J826 manikins. 3’

An example of the experimental input region within the Seat Factor Solver
spreadsheet is shown below in Table l. The first four columns (denoted by dark cells
with white font) are the only experimental inputs necessary for simulation. Deflections
of the H-pt. and knee relative to the position of the 2-D template [4] (dlh, dzk) were
chosen to have a positive down (into seat) coordinate system for Model IV simulations.

Positive relative angular deflections (Def Ang) are knee-elevating rotations.

l9

Table I. Input data acquired from experimental incremental loading of the manikin is shown in
white font within dark cells, along with calculated values based on experimental data shown in
grey.

H-PT KNEE

 

Step 2. Static Reaction Force and Moment Calculation

The static reaction forces and moments, which were based on experimental
loadings and measurements, were calculated for each incremental loading step according
to the equations of statics shown below. Figure 10 is repeated below as a free-body

diagram for statics-based analysis and derivation of equations 3 and 4.

20

'7 . . .
Slider prohibits horizontal translation of the l-I-pt.F

 

 

H-pt

 

Figure 10 repeated. Coupled force and moment kinetic model

R” = F” + FK (3)
M” =FK *LK *(cos(@m,c,,)) (4)

The set of reaction forces and moments varied slightly with the variation in angle
(9mm, ) between experimental results for each seat. These experimental reaction forces
and moments (equations 3 and 4 based upon experimental, incremental input forces F H
and FK) were then used as a basis of comparison to optimize the stifiiiess parameters
(K’s) within the modified Taylor series expansion matrices (equations 5 and 6) using the

Seat Factor Solver spreadsheet within MS Excel [3].

Step 3. Optimization of Modified Taylor Sena-Based Seat Factor Matrica
The modified Taylor series-based equations used to calculate the simulation
reaction force and moment based upon optimized seat factor parameters (K parameters)

were as follows:

F, =Z(,,Kfi, *5"+,,K,, *3") (5)
0

21

.9 *3") (6)

M, =Z(,,KM. *6"+,,K
0

The K parameters were Seat Factor stifiiress coefiicients used to create the
coupbd 2x2 matrices optimized in the Seat Factor Solver spreadsheet. These K
parameters are the calculated Seat Factors that define the mechanical interaction between

themanikinandtheseat.

A f 7', ‘ . , .
I ._
_)

<5

MS Excel Solver was used to reduce tthMS difference een the
experimental reaction forces and moments (equations 3 and 4) and simuhtion reaction
forces and moments (equations 5 and 6) by optimizing the K parameters within the
modified Taylor series-based, Seat Factor matrices. The linear model was optimized
first, and the K coeflicient results following optimization were used as a guess for the
first matrix in the second order expansion. The two matrices optimized for the second
order expansion were used as a guess for the first two terms of the third order expansion.
Using previous results as guesses for higher order expansions ensured proper
convergence (for all optimization procedures discussed within this thesis of models
higher than first order, guesses were used to ensure proper convergence to a solution or

set of optimized parameters).

The Seat Factor matrices optimization process could be done regardless of the
order of modified Taylor series expansion (number of 2x2 seat factor matrices). For this
project, models were generated using linear, second order, and third order modified
Taylor series expansions. As will be discussed in upcoming sections, increasing the

order of modified Taylor series expansion added terms to the simulation equations

22

(equations 5 and 6), which typically reduced the RMS difference between the calculated,

and simulation static reaction force and moment.

The number of sets of 2x2 matrices was dependent upon the order (n) of modified
Taylor series expansion i.e. Linear (first order): One, two by two matrix with a total of

form K Seat Factor parameters, Second order: Two, two by two matrices with four K

parameters per matrix, totaling eight K Seat Factor parameters, etc.

RMS error, or difference, was chosen for optimization procedures to minimize the
difi'erences throughout the data points as a mean, and to remove the effect of signs of the
differences. Optimization of the Seat Factor matrices was accomplished by calculating
theRMS difl’erence betweentheexperimentalreactionforcesandmomentsandthe
simulation reaction forces and moments. This number was then minimized, allowing the

K coeflicients of the Seat Factor matrices to change (optimize) using MS Excel Solver.

Step 4. Solving Equations to Determine Simulation Raids Based on Calculated Seat
Factor Matrica

\
‘.
\
l
4
i

._.—l

[1:0 solve for simulated deflections (5 ") and rotations (.9") using the optimized

set of Seat Factor matrices based upon equations 5 and 6, a MS Excel Solver optimization
process similar to that discussed in Step 3 was used. By minimizing the RMS difference
between the experimental and simulation static reaction forces and moments with a fixed
set of Seat Factor matrices and a set of H-pt. deflection and thigh rotation values to be '
optimized, the simulated H-pt. deflection and thigh rotation values could be solved within
a specific, RMS error range. In other words, this is the opposite procedure of Step 3 as
the Seat Factor Matrices were held constant to optimize the H-pt. deflection and thigh

rotation values, which solved equations 5 and 6 within a specific RMS error. From the

23

simulated H-pt. deflection and thigh rotation values, H-pt and m position could be
calculated to compare to the original set of experimental data. Based on numerical
analysis and graphical comparisons the models could then be validated within a known

error range.

Step 5. Graphical Analysis of Optimization and Simulation Raults

Within the SFS analysis worksheets, several plots were generated to display the
results of optimization and simulation of Seat Factor matrices. Graphical results for each
seat are shown in Appendix A. In the section below, select figures shown in each table

within Appendix A are explained.

Plots of H-pt. Force vs. H-pt. Deflection, and H-pt. Moment vs. Relative Thigh
Rotation, with both experimental and simulation paths present for comparison provided a
graphical method of error analysis for each order of modified Taylor series expansion.
Examples of these plots are shown for a third order simulation below in Figure 11a and
11b. Graphically, the differences between the experimental (points) and simulation (solid
line) responses could be analyzed. These plots also provided a graphical representation
of the non-linear (typically stifi‘ening) response of each seat. For most seats as seen in
Figure 11a, the seat stifi‘ened under the H-pt with increasing applied load. When knee
loads were applied following incremental H-pt. loading, which increased the total H-pt.
reaction force without adding additional load at the H-pt, the H-pt. position remained

stable or elevated (deflection decreased) slightly.

As seen in Figure 1 lb, incremental H-pt. loading with constant nominal knee

loading lead to an increase in relative rotation with constant reaction moment. Knee

24

loading with constant H-pt. loading increased the reaction moment and caused the

manikin to deflect into the seat and reduce the relative angle between the H-pt. and the

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

knee along the thigh.
300 Order Nbdel 3rd Order Nbdel
Hpt Reaction Force vs. Absolute Reaction Moment vs. Relative Rotation
Deflection
900 150000 J
800 ‘ 130000 c Mi(stat)
:33 . ’2‘ 110000
2 ,E 90000 “—M‘a'd)
V 500 z
8 400 : 70000
3 8
o Rh(stat) o
200 2 30000
100 — F(3rd) 10000 -
0 a , . _1 t 2 1 .
0 20 40 60 0000 0 2 4 6 8 10
Deflection (mm) Relative Rotation (deg)
Figure 11a Figure llb

Figure 11 a,b. Plots for comparison of experimental results (diamonds) to simulation results
based upon third order Taylor series expansion matrices (line) (Seat A).

Figure 12 is an example plot of thigh position as a fimction of H-pt and knee

loading for the third order model of Seat A. This plot was created for each seat to

graphically represent the deflection path of the linked H-pt and knee (analogous to a

human femur) during simulated incremental loading, and to determine the point of

rotation throughout H-pt. loading (similar to ETSL fi'om Models H and 111; see

conclusion 7). For most seats (excluding Seat C, a suspended mesh prototype), a finite

region of rotation during H-pt. loading was observable.

25

 

 

 

THIGH POSITION AS A FUNCTION OF H-PT AND KNEE LOADING
3RD ORDER MODEL

 

-200 Will: (N):
H-pL-Knee

 

 

 

+20
+185-29
+26429

 

 

 

 

 

 

 

 

 

 

 

Position Referenced to SAE Jszs Tern plate (nun)

 

 

O 100 200 300 400 500

 

Figure 12. Thigh position as a function of incremental H-pt. and knee loading. The graphical point
of thigh rotation (zero deflection) during H—pt. loading B approximately 225m from the H-pt. along
the thigh for this seat (Seat A).

Three-dimensional surface plots were generated to graphically display the H-pt.
reaction force and moment as a function of the H-pt. deflection and relative thigh
rotation. These plots are interpolations and extrapolations of the reaction forces and
moments beyond the path of experimental deflection and rotation based upon the
modified Taylor series expanded seat factor matrices for each seat (see Figure 13a,b).
Graphically, these plots provided individualized insight for each seat into the quality of
modified Taylor series expansion model-based interpolations and extrapolations within
and outside the range of experimental loading and deflection. Models that generated
regions of large, variant local curvature outside the range of experimental loading and
deflection had likely lost correlation to experimental results. Models with a continuous

trend are better for predicting manikin position outside the experimental path.

26

 

In most H-pt. reaction force surface plots, a slight to dramatic stifl‘ening trend was
observable within and around the range of experimental loading as the reaction force
typically increased with an increasing slope as H-pt. deflection increased. This is
expected due to th{stifl‘ening properties of seating fog Reaction moment response was
less consistent than the reaction force surface plots, and required individual analysis for

each seat.

27

 

RelativeThigh v r . ,
Rotation (699.) q. 53

  
  

FORCE SURFACE MODEL-3RD ORDER

H-pt Deflection (mm)

 

 

 

 

 

 

 

WSURFACEMGRDGIDER

Nag' Relaivenigh

Rm (deg)

 

 

 

Figure 13 a,b. Third order surface models of reaction force and moment as a function of H-pt.

deflection and relative thigh rotation (Seat A).

28

 

 

Step 6. Occupant Position Appron'mations

Using average loads based upon a large set of human subject data gathered in

dissertation work by Bush [7], sets of typical reaction forces and moments were

calculated for small female, medium male, and large male body types. Using these loads

as inputs into the Individual Deflection-Rotation Solver module in the Seat Factor Solver,

third order Seat Factor Matrices interpolated and extrapolated mathematically to solve for

the approximate H-pt. deflection and thigh rotation due to these particular imposed loads.

Since the thigh surface contours of the SAE J826 and ASPECT manikins are intended to

represent 50% males (body dimensions and deflected contours), results for a small female

or large male based upon SAE 1826 or ASPECT manikin experimental data acquisition

are only estimated. This module of the Seat Factor Solver is shown below in Table 2.

Table 2. Individual Deflection-Rotation Solver for Occupant Position Approximations.

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

HmflolvaeLawF __ . .. flfidOOCLPANTPosrnmAPI-‘RCNMAHOBS
“MIOW'I—CW .V - . 2550mm FH FT F M

|sr= 271 112 am 25378
[MM 43 129 537 23230
ILM 4a 135 as 30817
_ worms :2le 051mg
INDIVIDUAL DEFLECTION-RO'I'ATION SOLVER SF 358 4.9
F M dzh DerArg Enu’? RSE w 456 68
25373 3581 4.9, 1. 000m * um 47.2 7.0

 

 

 

 

 

The terms F n and F1 within the table above were the average measured forces in

Newtons under the H-pt. and thigh (235 mm forward of the H-pt.) respectably, for each

of the three body types observed in experimental data acquisition by Bush [7]. The terms

F and M were the true experimental reaction force and moment components in Newtons

29

 

based upon F H and F1, which were used as inputs into third order Model IV Seat Factor
matrices to calculate the H-pt. deflection (dzh) and the relative thigh rotation (Def Ang).
In the lower right comer of the table, results fiom average loads for three difl‘erent body

typesarelisted.

This module proved the ability of the Seat Factor matrices to interpolate and
extrapolate results outside the path of experimental loading. Though only approximate,
this module provided comparative insight into the deflected positions of the manikin
when loading criteria indicative of transmitted loads outside the experimental loading
path for three different body types were used as inputs. As expected, loads indicative of a
small female deflected the manikin into the seat less than loads indicative of a medium
male, which deflected the manikin into the seat less than loads indicative of a large male.
Results for each set of deflections and rotations due to loads indicative of these three

body types are shown in Appendix A for each of the 30 seats analyzed within this project.

30

Results and Conclusions from Model IV

Model IV and corresponding mathematical methods are an effective way to
determine quantitative Seat Factors that define the static interactions between seating
manikins and automotive seat pans. The ability to generate equations that define the
structural properties of seats as a system of multiple components based upon
experimental data acquisition provides insight into the individualized behaviors of seats
when subjected to varying loads. The following statements are conclusions based upon

mathematical modeling analysis during the course of this project:

1. Simulation error for Model IV is significantly less than the current H-pt.
positioning tolerances and less than the certainty of data acquisition. Current
automotive H-pt. positioning tolerances allow error in H-pt. positioning of up to
20mm during occupant positioning and design procedures. As observed in Table 3
below, Model IV defined the structural properties of seats and predicted deflections
and rotations within a mean error much less than 20mm for each of three orders of

modified Taylor series expansion models tested.

As shown below in Table 3, Model IV had a mean error less than that of
experimental data acquisition certainty at the second and third order modified Taylor
series expansion level (data acquisition has an uncertainty of approximately i2mm).
The ability to model and predict the static deflected position of a manikin’s buttocks
and thigh region based upon experimental data acquired fiom industry standard
manikins is a significant advancement in the study of biomehanics and occupant

positioning.

3]

Table 3. Average H-pt. deflection, knee deflection and relative thigh rotation difference across
experimental loading range for linear, second, and third order Seat Factor Matrices.

VG. DIFFERENCE BETWEEN EXPDATA AND SIMULATION ALONG EXP. LOADING PATH

1.

1 .75
0.67
2.37
1 .82
0.74
2.40
2.1 1
0.88
1.1
2.35
4.07
1.74
1.75
1 .68
1.04
1

d
d
d ‘

I0
d
ooopppopooo
b
«at

0b

are
pooopoo
.3

to
..n

1
1.
1.
O.
6
1
1
O
O
1
1
1
1
1
O.
1
0
O.
1
0
0.

7 0.63

 

' F is a statistical outlier due to convergence error in SFS. and therfore is not included in statistical calculations

2. Calculated Seat Factor equations can be used to interpolate or extrapolate
results outside the path of experimental data acquisition. Using a set of H-pt.
deflections and thigh rotations within a realistic range of deflection as inputs, the Seat

Factor matrices can be used to interpolate or extrapolate reaction force and moment

32

results outside the path of experimental data acquisition. These data sets were
displayed as three-dimensional surface plots in Figure 13a,b, and for each seat within
Appendix A. An example of this interpolated and extrapolated surface with
experimental path overlaid on the surface is shown below in Figure 14. The
experimental path appears to be segmented since it weaves above and below the
suriace generated by the Seat Factor matrices. This is due to the slightly inexact
surfiwe fit of the modified Taylor series expanded, Seat Factor Matrices to the line of

experimental data acquisition

 

 

 

 

H-pt. Deflection (mm)

D -5 Relative Thigh Rotation (deg)

 

 

Figure 14. Interpolated and extrapolated H-pt. force surface Plot, shown with extrapolation and
interpolation beyond the path of experimental data (Seat 0).

The coupled set of equations (equations 5 and 6) could also be solved for H-pt

deflection and thigh rotation results based upon any input force and moment system.

33

This was described within Step 6 of the Mathematical Modeling Methods section of
this thesis. Simple static equations (equations 3 and 4) are necessary to determine
input reaction force and moment based upon a system of forces at the H-pt. and
another point along the thigh. For example, Bush measured forces under the buttocks
region corresponding to the H-point, and at a region centered 235 mm forward of the
H-pt. along the thigh [7]. These measurements can easily be transformed into an

equivalent reaction force and moment set at and around the H-point.

. Interpolation and Extrapolation accuracy may decrease for loading inputs that
are different from the load/deflection path of experimental data acquisition. For
this reason, it is important that the path of experimental data acquisition extends
beyond the range of plausible simulation loading. This means that if simulation
results and calculated Seat Factors rmtrices are required for a particular body type,
the experimental loading protocol should be similar or preferably span below and
above those loads necessary for simulation. Therefore, interpolation of small fermle
or medium male results using seat factor matrices calculated from a large male
experimental loading procedure should produce more accurate deflection results than
extrapolation of small female experimental loading to a large male simulation.
Another fundamental reason for acquiring data points beyond the range necessary for
simulation is that the Taylor series expansion results can drift by loss of correlation at
the extreme experimental data points at each end of the experimental loading and

deflection path.

34

4. H-pt. deflection and thigh rotation can be referenced to any reference position.
Data acquired for this project used the SAE J826 [4] two dimensionaL plastic,
drafling template as the initial location to which all other manikin deflections were
referenced i.e. relative thigh rotation. Within Model IV, H-pt. deflection and thigh
rotation were relative measures that were referenced to a set of initial points. This
procedure was chosen to attempt to locate the position of the manikin with respect to
an approximately “weightless” buttocks-thigh contour in contact with the seating
surface. Though this was an effective procedure, H-pt. deflection and thigh rotation
could be referenced to any known reference position that would be convenient for
analysis with modifications to the Seat Factor Solver spreadsheet. Alternative
reference positions include the loaded SAE J826 manikin location as prescribed by
the SAE 1826 H-pt. procedure [4], or the loaded ASPECT manikin position as
establismd by the ASPECT program [5]. The two-dimensional contour of the SAE

J826 2-D template [2] is shown below in Figure 15.

NNPOINT

 

TRAVEL PATH

Figure 15. Two-dimensional contour of the SAE‘J826 Manikin [4]. A V4 inch, 2-D plastic
template with rods extending from the H-pt. was used as a reference for data acquisition by
Hubbard and Gedratis [l] and Hubbard and Radcliffe [6] within this project [4].

35

5. Adding additional orders of modified Taylor series expansion to the set of Seat
Factor matrices reduced the RMS error, but increased the set of possible
solutions. Each additional term added to a Taylor series expansion firrther reduced
the error in approximating the original function, but further increased the set of
possible solutions. This was observed in the difference in mean RMS error and mean
percent error as shown at the bottom of Table 4 below. A diminishing decrease in
difference occurred with the addition of each order of Taylor series expansion to the
set of Seat Factor Matrices i.e. the reduction in difference was greater between the

linear and second order model than between the second and third order model.

The MS Excel solver tool is iterative and uses an initial set of numbers to guess at
a solution based upon user prescribed criteria. To ensure a coherent result, a
reasonable solution must be used as a starting point for iteration. The larger the order
of Taylor series expansion, the better the guess must be to allow for iteration to a
logical solution set. In order to initially solve for the seat factor matrices, the linear
matrix ~should be calculated first, and then used as a guess for the first order matrix in
the second order Taylor series expansion. For each additional order of expansion (71),

the n-l order expansion should be used for the first n-l matrices.

In order to solve for the results of the simulation, a good data set to use as a guess
for the solution is the experimental result. From this solution, the solver will iterate

to a solution based upon the chosen calculated seat factor matrices.

36

Table 4. RMS and Percent Error vs. Order of Taylor Series Expansion.

VI.

1

 

*FisastfisficflafliadatomgamminSFS, atdfltafueisruitdtxbdinsuisfidm
. Order of Taylor series expansion should be minimized based upon error
tolerances. As previously discussed, the addition of Taylor series terms to the model
decreased the error, but added to the overall complexity of the seat factor set. Based
upon the error tolerance of the application, the lowest possible order of modified

Taylor series expansion should be used. This will reduce the number of possible

solutions, and reduce the overall complexity of the model. As user application

37

tolerances are not curremly known for each application, suggesting a specific order of
expansion as the best model would be an arbitrary statement. The best choice for
expansion may vary with each seat modeled based upon the type of response and

degree of non-linearity during experimental data acquisition.

. The thigh lines of the manikin in all seats pivoted about a specific region of the
seat during H-pt. loading with constant applied knee load. As a fundamental part
of the meoretical basis of Models H and III, the link between the H-pt and the knee
can rotate about a small region or point between the H—pt. and knee during H-pt.
loading with constant kme loads. This is described for these models by the ETSL
term (Equation 1). For Model IV, this assumption was verified as the phenomenon
was visible using graphical analysis (see Figure 12) of the simulation results based
upon a third order expansion. For each seat, it was apparent that there was an
approximate point (region) of zero deflection or thigh rotation. This point is the
center of thigh support during H-pt. loading with constant, nominal knee loads. This
point likely changes position during knee loading of the manikin in the seat. Future
work could determine whether there is a comfort correlation between subjective

testing and the position of the thigh support.

38

CONCLUSION

The objective of this research project was to develop models that described and
simulated the force-deflection response of the buttocks-thigh region of industry standard
seating manikins in interaction with automotive seats. This objective was met by
successively developing and evaluating several kinetic models using validation
techniques with 30 automotive seats/conditions. Models I-III proved to be inadequate in
modeling the force-deflection/rotation response using physical modeling techniques. At
the conclusion of this project, Model IV in conjunction with the Seat Factor Solver
spreadsheet in MS Excel [3] proved to be the best model and method combination for

describing and simulating the response of seating manikins in automotive seats.

Model IV used modified Taylor series expansions to describe and simulate the
force-deflection/rotation behavior of the buttocks-thigh region of manikins as observed in
experimental testing. The ability of Model IV to model the force-deflection/rotation
response of the manikin in an automotive seat was proven to have an error at the second
and third order level of modified Taylor series expansion less than the uncertainty of
experimental data acquisition. This constitutes extremely successful validation of this

kinetic model.

Model IV and corresponding mathematical modeling techniques could be very
useful for software models of occupant positioning in automotive interiors. Model IV
my also be of interest to occupant positioning software models for aeronautical, office,
or heavy machinery applications. Currently, software models assume the surface of the
seat to be rigid, and do not account for the properties of the foam and suspension within

individual seats. Model IV and corresponding experimental data acquisition and

39

mathematical modeling methods could dramatically improve the ability of a software

model to position an occupant correctly within a seat and interiorjml

Future Work

Though this work is a major step in developing models and methods for defining
the static characteristics of automotive seat pans with respect to a progressively loaded
manikin, thisproject isonlyafirst stepwithinthe fiillscopeofkineticmodeling of
manikins and occupants in automotive seats. As with most initial investigations, analysis
begins for the simplest scenarioEthis case, static deformation of an automotive seat
due to imposed loads by a seating manikin was investigated, and subsequent seat factor
matrices were developed and calculated to define the mohanical characteristics of each
seat. The next step is to develop dynamic models and experimental data acquisition

techniques that investigate and define dynamic characteristics of seats when dynamic

A

loads are applied indicative of an uneven road surface. This is an important topic, as K

seats are an integral component of interaction between the human occupant and the
vehicle, and ultimately the road surface traveled. Suspension characteristics of
automotive seats may be of similar importance as the vehicle’s main suspension system

”"1
. . . . l
in terms of the overall user’s impressron of the vehicle.

_,__J

One of the important advantages of the mathematical modeling technique used in
conjunction with Model IV is the ability of the Seat Factor Matrices to predict force-
deflection/rotation response outside the range of experimental loading. To validate this ‘ “v -.
theory, testing could be conducted which used an initial data set to model and predict the if i: I I
response of the seat to a manikin loaded outside the range of experimental loading.

Subsequently, an experimental loading procedure that investigated the same loading as

40

checked in simulation outside of the experimental loading path could be pursued to
validate the ability of the model to predict position of the buttocks-thigh region of the

manikin.

Another topic of great importance, is determining possible correlations between

Seat Factors and subjective comfort ratings of individual automotive seats. This requires

_.4 -w-r—y-

comparative analysis of Seat Factor sets as compared to subjective responses to questions

regarding individual seat comfort in various positions and regions relating to each seat.£lf L.

correlations between subjective comfort ratings and quantitative Seat Factor sets could be
established, the seating design process would be aided dramaticallijhis work my be

of interest to seating manufacturers in automotive and other industries.

This project investigated the interaction between the H-pt. to knee link and the

automotive seat pan. Though this is an important interaction interms of occupant

positioning and possibly comfort, many other interactions between the occupant/manikin I '

and automotive interior could be modeled. One proposition is to extend the current
Model IV analysis techniques to additional segments of the occupant/manikin in a linked

system. This model may appear as shown below in Figure 16.

4]

fi«

I
f

. l/ '7

FN

 

 

O
Fr
F L
w i 0“} MK
3 ..
FP -333%
MP i Mu
FK . O
F“ :3
.w MF
Fr

Figure 16. Proposed two-dimensional linked system of reaction forces and
moments (dark arrows) with input loading (white arrows) at crucial joints.

The ability to model the complete interaction of an occupant/manikin and an automotive

seat/interior is contingent upon the ability of researchers to experimentally measure

transmitted forces and specific deflected locations at crucial joints using a seating

manikin.EheASPECT manikin [5] has an articulating back section with jointed regions

similar to the kinetic model proposed in Figure l6; If incremental transmitted loads and
_..t

manikin joint positions could be measured experimentally, this type of kinetic model may

be feasible for modeling, simulating, and gaining predictive postural positioning

42

knowledge for a two-dimensional occupant in an automotive interior. The ability to
describe and predict the posture of a complete occupant in an automotive interior due to
variable inputs is the ultimate objective for kinetic modeling of manikin-automotive

interior interaction

43

APPENDICES

APPENDIX A

45

APPENDIX A

The manufacturers and types of seats corresponding to letters used within this

document are provided as a legend in Table L-l.

Tables A—l through A-30 represent the analysis of 30 seats/conditions using
Model IV and the Seat Factor Solver spreadsheet. Tables A-l through A-22 represent
data acquired by Hubbard and Gedraitis [1]. Tables A-23 through A-3O represent data
acquired by Hubbard and Radcliffe [6]. Tables A-23 through A-26 were tested using the
Prototype ASPECT manikin. Tables A-27 through A-30 were tested with the SAE J826
manikin for comparison to Tables A-23 through A-26. Within each table, the seats are
analyzed mathematically and then graphically to describe their response to manikin
loading. Colunms of analysis for 3rd order, 2"‘1 order, and Linear models show calculated
Seat Factor matrices and the subsequent RMS difference associated with each modified
Taylor series expansion. The Seat Factor matrices based upon equation 5 and 6 are

labeled as Stifliless Matrices, and are transposed within the spreadsheet to ease display.
Next to the transposed Seat Factor matrices, the variable ‘x’ refers to 6 " and ‘ang’ refers
to .9" in equations 5 and 6. The ‘Significance Magnification on M Calculation’ was

always one, or equal weighting on optimizing the F and M components according to the

calculation of the RMS difference between experiment and simulation.

For seats 0 through 9 or Table A-15 through A-22, the raw data from acquisition
by Hubbard and Gedraitis [l] for the first 2-D template reference position was lost. As

this was a predominantly a validation procedure, 13.64 mm was added to the H-pt., and

46

6.44 m added to the knee position of the unloaded manikin (step 2) to establish a

reference position for modeling these seats. These values are the average differences

betweenthe step 1 and step 2 positions measured by Hubbard and Gedraitis [1].

Table L-l. Seat Key for seats analyzed within the Seat Factor Solver Spreadsheet.

 

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BMW/Renee

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APPENDIX B

168

APPENDIX B

Appendix B contains an umnodified comparative report of results fi'om testing
four seats with both the SAE 1826 [4] and ASPECT [5] manikins, entitled: ‘Seat pan
load-deflection measurements for determination of seat factors with the ASPECT butt-
thigh (ABT) segment of the ASPECT manikin prototype at MSU and the SAE J826
manikin’ by Robert Hubbard and Nathan Radcliffe. This original report is presented as
supplemental material with description of experimental methods and results not fully
discussed within the body of this thesis. References made within this report refer to the

reference section of Appendix B, and not the body of this thesis.

169

Seat pan load-deflection measurements
for determination of seat factors with
the ASPECT butt-thigh (ABT) segment of the ASPECT manikin prototype at MSU
and the SAE J826 manikin

Robert Hubbard and Nathan Radcliffe

A. Introduction:

Seat factors were first measured and documented by Hubbard and Gedraitis [ l] in the development
program for the ASPECT manikin [2]. This initial measurement of seat factors was done using the SAE
J826 manikin [3]. In a recent study for her BS. in engineering at Michigan State University (MSU),
Bogard measured the position of the H-point and the lumbar prominence using the new ASPECT manikin
[4]. Bogard’s measurements were done in seats that were used by Gutowski [5] in posture measurements
with human subjects.

The purpose of the present study is to measure the vertical deflection response of the ASPECT manikin
butt-thigh segment (ABT) and the SAE J826 as they are progressively loaded into selected automotive
seats. The vertical load-deflection measurements are the basis for interpretations of seat pan responses,
which are important seat factors

B. Methods:

3.1. Loafing of the ASPECT butt-thigh segment (ABT):

As in our previous work [1], weights were progressively added to the ABT of the ASEPCT manikin
prototype that we have at MSU. As the weights were added, the heights of the H-point axis and ham of the
ABT shell were measured above the floor. The H—point axis was measured at the same distance from the
center on both the right and lefi sides of the ABT. The height of the front of the ABT was measured at the
upper edge of the ABT shell on the mid-line. Weights were added first to the rear part of the ABT on or
near the H-point axis, and then weights were added toward the front of the ABT.

So that the loading of the ABT could be known In terms of equivalent loads at the H-point axis and a lateral
axis through the front of the ABT the ABT was suspended from lead cells that were attached to these two

axes. We recorded as weigWsuelyaddeQ The lateral axis

through the front of the ABT was 384 mm in front of the H-pomt ax15. \
cords to front of ABT

0

 

Figure l. Weighting of the ASPECT butt-thigh segment (ABT)

 

Figure 1 shows the ABT hanging by cords attached to H-point axis and a lateral axis through the fi-ont of
the ABT. Threaded steel rods (5/16 inch diameter, 5 inches long) were used to extend the HJC axis of the
ABT. Figure 1 also shows the ABT with all of the weights added in the loading steps shown and listed in

170

Table l. The torso weights with the ASPECT manikin prototype at Michigan State University (MSU) are
built up with eitha‘ six or five steel plates. The weights of the MSU ASPECT manikin are partitioned
differently than the final and standard ASPECT manikins, but the weighting of these manikins lS

equivalent. (Ind step 0 in Table l was accomplished by placingthe 1826 drafting template in the fest; D, ‘

The H-point axis for the .1826 template was extended to both sides with a rod so that the height of the H-
point could be measured while the J826 template was in a seat.

as shown in Figure l and Table l. ues of moment about the l-l-point axis in Table 2 were computed
as the load measured at the front 0 e ABT multiplied by the distance from the Hopoint axis to the front of
the ABT. In actual seats, this distance is at a small angle (typically less than ten degrees) and the moment
is computed with the cosine of that angle.

Table 2 gives the loading of the Hflmtaxis and moment about the H-point axis for the ABT and weights
Th

The H-point of the J826 template is 100 mm above the bottom of the template, which is the same distance
M the H-point axis to bottom of the ABT directly under the H-point axis. However, the contour of the
ABT under the H-point is different fi'om the J826. The ABT contour was based on measured human
contours and the region around the bottom of the pelvis is more prominent and less flat than the 1826. This
is probably the reason that the H-point axis with the ABT typically is approximately 10 mm lower in

seat than the with J826 once both manikins are loaded into seats, which is apparent in the results belo‘al

Table 1. Weight placements for loadingstemfor the MSU ASEPCT manikin butt-thigggment.

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

Load step q

0 J826 template for zero load reference (a

1 No weights added to ASPECT butt-thi ent

2 Two 6-plate H-point weights on H-point axis

3 Step 2 plus two S-plate H-point weights on H-point axis

4 Step 3 plus two S-plate Hyoint weights on H-point axis

5 Step 4 plus two torso weights on H-point axis just inside shell

6 Step 5 plus two torso weights on H-point axis just inside shell with slot down and back rear
edge restinflgainst shell

7 Step 6 plus two torso weights behind center structure with slots forward

8 Step7 plus two torso weights behind center structure with slots forward just outside weights

added in 7

9 Step 8 plus two torso weights on center structure with slots up and back lower edge against
nylon busthgOAm two torso weights were not used)

10 Step 9 plus two thigh wei

11 Step 10 plus two thigh weights

12 Step 1] plus two thigh weights

I71

T the

    

Step Load on Front Moment
H-poin about

Figure 2. ASPECT manikin butt-thigh segment in a
Saturn seat with all of the weights positioned as

the ABT for the loading steps of Table l.

  
   

described in Table land as shown in Figure l.

 

13.2. Deflections into seats of the ABT when loaded

Figure 2 shows the ABT in a Saturn seat with all of the weights positioned as described in Table 1 and also
shown in Figure 1. Tape was added to the weights on the thighs to keep them from sliding rearward. The
positions of the H-point axis and hum of the ABT were measured up from the floor using a meter scale

with millimeter graduations.

3.3. Loading of the J826 manikin:

Seat pan stiffnesses were initially measured in our previous work
using the SAE 1826 manikin [1,3]. Measurements using the 1826
manikin in selected seats were repeated in the present study for
comparison with similar measurements using the ASPECT
manikin as described above. The back angle of the 1826 manikin
was adjusted to be vertical for each loading set.

Figure 3 shows the 1826 manikin is a Chrysler LH seat loaded
with all of the weights in it as described in Tables 3 and 4. The
loading used in the present study was the same as in our previous
study [3] except for the loading of the knee axis. The loading of
the 1826 manikin was similar to, but not the same as, the loading
with the ASPECT manikin as described in the previous sections,
8.1 and 8.2. As with the ABT the cosine of the thigh angle was
not used to compute the moment in Table 4 but was used in data
analysis.

In the results section below, the responses Earn the ASPECT
manikin and 1826 manikin will be compared on the basis of load-
deflection of the H-point axis and moment-angle about the H-
point axis.

    

Figure 3. SAE .1826 manikin
in a Chrysler LH seat with
all of the weights positioned
as described in Table 3.

 

Table 3. Weight placements for loading steps with the SAE 1826 manikin.

 

Load step

 

1826 template for zero load reference

 

No weights added to SAE 1826 manikin with torso segment vertical

 

Two disk torso weights on H-point axis

 

Step 2 plus two disk torso weights on H-point axis

 

Step 3 plus two disk torso weights on H-point axis

 

Steg4plus two disk torso weights on H-point axis (Two 1826 torso weights were not used)

 

Step 5 plus two cylinder thigh weights

 

Step 6 plus 10 pormds (44.5 Newtons) hung from knee axis

 

 

”\IGMbWN-‘O

 

Step 7 plus 10 pounds (44.5 Newtons) hung from knee axis

 

Table 4. Loads at the H-point and knee axes of the
SAE 1826 manikin for the loading steps of Table 3.

 

 

 

 

 

 

 

 

 

 

 

 

 

Load
Step Load on Load on Moment
H-point Knee about H-
ax'n axis point
Newtons Newtons N-mm
0 0 0 0
l 170 27 12300
2 245 28 12800
3 322 28 12800
4 395 32 14600
5 474 31 14200
6 501 70 32000
7 505 113 51600
8 506 158 72200

 

 

 

 

173

 

C. Results and Discussion:
C1. Measurements of deflections due to seat pan loading with the ASPECT butt-thigh segment
(ABT)-
Table 5 is the raw data for the deflection measuranents of ABT into four automotive seats: Chrysler L11
(L11), Neon, Saturn, and Chrysler Town and Country (TC). Height measurements were made relative to

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

the flat, level floor.
Similarly, Table 6 is the raw data for the deflection measurements of the SAE 1826 manikin in these same
four automotive seats.
Table 5. Raw data for height measurements of ABT H-point axis and front of ABT
hip hip front hip Hip from hip hip front hip hip front
Load
Step LH LH LH Neon Neon Neon Satrn Satrn Satrn TC TC TC
right left ri t lefi Right left right lefi
mm Mm mm mm Mm mm mm mm mm mm mm mm
0 445 445 51 l 437 459 557 410 410 521 465 468 577
l 448 446 476 448 445 515 416 412 481 463 462 525
2 434 429 482 436 431 519 406 403 491 454 452 533
3 422 417 488 426 417 530 397 394 494 448 447 536
4 412 411 492 412 406 536 392 393 502 443 437 541
5 405 410 499 407 401 542 384 381 510 441 434 541
6 400 407 501 402 396 545 379 378 515 439 429 541
7 399 399 503 400 394 551 378 373 525 435 427 542
8 394 393 506 396 388 560 375 371 525 432 422 545
9 392 391 510 394 387 560 373 371 525 427 418 547
10 390 390 502 393 388 545 373 371 514 424 417 536
l 1 391 391 498 391 382 538 370 367 501 423 413 528
12 388 391 494 388 383 529 370 367 499 420 412 521
Table 6. Raw data for height measurements of SAE 1826 H-point axis and knee axis
hip hip knee knee hip hip knee knee hip hip knee knee hip hip knee knee
Load
L11 L1! L11 L11 Neon Neon Neon Neon Satan Satrn Satra Satrn TC TC TC TC
right left right left right nen right left right no right left right 1‘“ right k“
mm mm mm mm mm mm mm mm mm mm mm mm mm mm mm mm
0 445 445 511 511 456 441 556 556 407 411 521 521 460 462 573 573
1 430 431 517 515 429 427 559 556 400 396 522 519 456 450 561 555
2 425 429 520 519 421 415 566 560 395 390 532 527 455 442 562 554
3 419 420 526 526 413 412 568 566 388 386 535 530 441 438 564 562
4 410 414 536 532 408 403 574 570 383 381 536 532 430 426 567 563
5 403 407 538 537 402 397 573 570 381 377 537 533 419 418 570 569
6 404 402 530 529 400 396 566 565 379 375 532 527 415 413 558 554
7 403 403 521 518 400 397 554 551 380 375 523 517 414 415 538 534
8 403 403 503 503 399 395 534 530 377 374 504 499 415 416 518 516

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

174

 

 

For the measurements using the ASPECT manikin, the seat pan loading of Table 2 can be combined with
seat pan deflection measurements of Table 5 to provide load0deflection responses. Similarly, with
measurements using the SAE 1826 manikin, the seat pan loading of Table 4 can be combined with seat pan
deflection measurements of Table 6 to provide load-deflection responses. Radcliffe [6] is working to
interpret these measurements in terms of:

1. vertical translational deflections of the H—point axis in response to downward loading tlnrough the H-

point axis and

2. rotational deflections of the line fi'om the H—point axis through the knee axis.
By fitting these translational and rotational respornses with a mathematical model, the response of a seat pan
can be characterized so that loadings other that those used for the measurements can be predicted. Once
these models are available, they will serve as Seat Factors.

Figures 4 though 7 show the measured responses of fan automotive seats (Chrysler LH, Chrysler Neon,
GM Satin-n, and Chrysler Town and Country) to loading of the 1826 and ASPECT manikins referernced to
the 1826 template. For each seat, the first plot (a) shows the force-deflection response of the H-point axis;
the secornd plot (b) is the moment-angle plot relative to the H-point axis. Notice that the vertical scales are
difl‘erent for the difl’ercnt seats.

 

H-pt Force vs. H—pt Deflection

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

AapectvaABJMSeatingManifina
Fifi-mm
-10 6%
E O
z 10-
§ 2.
. +LHIBI'
30
+4.11.”
40
s
x 50
60

 

 

 

Figure 4a. Force—deflection responses of a Chrysler LH seat with the ASPECT (ABT) and 1826 manikins

175

 

H-pt. WtuRelativeThighRotation

 

 

Aapect vs. SAE 1826 Seating Manikin
H-pt. Ital-mm
4.00
4.00
-2.00

 

 

0 20000 30000 40000 50000 60000 70000 80000

4

 
 

 

2.00 .

4.00

0.00 )

8.00 (fa/K . +01.“
10.00

12.00

 

 

 

 

 

 

 

RalaIva Tlnlfln Won (0 Race lav“
Weaning.)

 

 

 

 

 

 

Figure 4b. Thigh moment-angle respornses of a Chrysler LH seat with the ASPECT (ABT) and 1826
manikins

176

 

H-ptFoscevaH-ptbeflection

massajmmm
Marco-m
o 100 zoo 300

400 500 000

 

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O

 

 

 

 

 

 

 

 

:20
.2,30

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340
‘50 +n~m~um
‘ao

 

 

 

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O

 

 

 

Figure 5a. Force-deflection responses of a Chrysler Neorn seat with the ASPECT (ABT) and 1826 manikins

 

H-peMamnnuRelativeThighRotaflon
AspeavaSABJMSeatingManikina

HUMM

10.0.».3!’
88888

as».
88

10.00
12.00
14.00

Relative Thlgh Rotation (0 Knee lavatlna
Rotation) (a...)
a.
8

 

 

 

 

Figure 5b. Thigh moment-angle responses of a Chrysler Neorn seat with the ASPECT (ABT) and 1826
manikins .

177

 

3.9112”tuan

AapectvaSAEJMSeatingManildna
MM“)
-10

 

100 200 300 400 500 0

\ +smasr
iffy: -....-

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O

 

 

 

 

8

ma M «mum
B

 

 

 

8

 

 

 

 

 

 

Fignne 6a. Force-deflection respornses of a GM Saturn with the ASPECT (ABT) and 1826 manikins

 

H—ptMomentvaRelativeThighRontion
AapectveSAEJWSeatingManiltine

III-n. worm)

 

 

 

 

 

 

 

 

 

 

 

 

 

 

Memmnommm
Iatatloanag.)
A N
8 8

10.00
12.00 I

 

 

 

 

 

Figure 6b. Thigh moment-angle responses of a GM Saturn seat with the ASPECT (ABT) and 1826
manikins

178

 

H-pt PorcevnH-ptDeflcction
AspectvaSAEJMSeating Manflcina

"ii-ml")
0 100 200 300 41!) 500 600

x L

 

O

 

/

 

 

\ -.....
\ -..—....

 

 

 

 

 

 

“Mlomlm
8 8 8 8 8

 

 

 

 

Figure 7a. Force-deflection responses of a Chrysler Town and Country seat with the ASPECT (ABT) and
1826 manikins

 

H-fihionentnkelative'l'highkotation
AapectuSAEJMSeatingManildna

”m0?“

 

 

 

 

 

.0 .N p .I'o .1. .6»
88 8 8 8 8
H-e

 

 

Relative midi Rotation (0 Knee Mating
Rotation) (deg)

 

 

SIX)

 

 

 

Figure 7b. Thigh moment-angle responses of a Chrysler Town and Country seat with the ASPECT (ABT)
and 1826 manikins

The force-deflection plots (a) are all similar. With the 1826 manikin, the respornses are nearly linear. With
the ASPECT manikin, there typically is an initial offset, tlnen a region of softer respornse from
approximately 100 to 200 Newtons, then the stifliness with the ASPECT and 1826 manikins are nearly the
same from 300 to 400 Newtons. The initial offset and softer response with the ASPECT manikin is
probably due to the diflerence in contour shape under the H-point axis. As mentioned above, the ASPECT
manikin is more prominernt due to realistic pelvis shape of tlne human data.

If a single seat pan stifliness were to be selected as a Seat Factor, we suggest using the stifiness measured
with the ASPECT manikin between 300 and 400 Newtons on the H-point axis. However, Radcliffe [6] is

179

developing a mathematical modeling process that represents the complex and nonlinear responses of
automotive seats to loading with both the ASPECT and 1826 manikins.

The moment-angle plots (b) are also similar. Starting fiom the origin, which is the orientation of the 1826
template, the knee axis of the 1826 manikin (without added weights) was typically 2 degrees higher than
the 1826 tennplate. The exception was with the Chrysler Town and Conmtry seat where the 1826 manikin
knees wee slightly [owe than the tenplate. The Town and Conmtry seat had a crease in the midlinc of tlne
seat pan, which was a problem for placement of the of tlne J826 template.

As the H-point axis of the 1826 manikin was loaded without sigrnificant change in moment about the H-
point axis, the lance axis of the 1826 manikin rotated upward a few degrees. After the initial loading of tlne
H-point axis and with an increase in moment about the H-point axis, the lance axis rotated downward into
the seat.

With the ASPECT butt and thigh segment (ABT) in each seat, the float of the ABT was approximately 4
degrees lower than the knees of the 1826 manikin. As the H-point axis of the ABT was loaded without
increased moment about the H-point axis, the front of the ABT rose and this upward rotation of the ABT
was consistently greater than the corresponding rotation of the 1826 manikin. Again, tlnis difference is
probably due to the more prominent contour of the ABT under the H-point axis; with H-point loading, the
ABT moved down in the seats with H-point and the thighs rotated up more than with the 1826 manikin,
which has a flatter contour than the ABT.

With increasing moment about the H-point axes, the moment-angle responses of the seats were equivalent
with both the ABT and 1826 manikin. For a single parameter to represent support of tlne thighs by a seat,
we suggest the slopes of the moment-angle responses between 20,000 and 40,000 Newton-millimeters.
However, as stated above, Radclifle [6] is developing a mathematical modeling process that represents the
complex and nonlinear responses of automotive seats to loading with both the ASPECT and 1826 manikins.
We believe that this development by Radclifl‘e is preferable to using simpler seat factors, such as force-
deflection and moment-angle stimnesses.

References:

l. Hubbard, R., and C. Gedraitis, “Initial Measurenents and Interpretatiorns of Seat Factors for the
ASPECT Program”, Soc. of Auto. Engin. (SAE) Paper Number 1999-01-0958, presented at the 1999 SAE -
International Corngress, March, 1999.

2. Schneider, L., M. Reed, R. Roe, M. Manary, C. Flannagan, R. Hubbard, and G. Rnnpp, “ASPECT: The
Next Generation H-point Machine and Related Seat Desigrn and Measurement Tools”, Soc. of Auto. Engin.
(SAE) Paper Number 1999-01-0962, presented at the 1999 SAE lntenational Congress, March, 1999.

3. “Devices for Use in Defining and Measuring Vehicle Seating Accommodation”, Society of Automotive
Engineers Recommended Practice 1826.

4. Bogard, H., “Additiornal Methods to Characterize Lumbar Supports Using the ASPECT Manikin”, thesis
for the Bachelor of Science Degree, Department of Materials Science and Mechanies, Michigan State
University, 2000.

5. Gutowski, P., “Influence of Automotive Scat Factors on Posture and Applicability to Designn Models”,
thesis for the Master of Science Degree, Department of Materials Science and Mechanies, Michigan State
University, 2000.

6. Radcliffe, N., thesis in preparatiorn for the Master of Science Degree, Department of Materials Science
and Mechanics, Michigan State University, 2001.

180

APPENDIX C

181

APPENDIX C

This section describes and derives a secondary method for determining the ETSS
in Models II and 111, which was not investigated fully due to the efi‘ectiveness of Model

IV in modeling the interaction between seating manikins and automotive seats.

 

 

_ Initial Position
— Positiorn alter H-pt. Loading

Figure C-l. Descriptive two-dimensional figure of H—pt-knee deflection and
rotation for analysis of the ETSS.

ZF,‘=0=F,, +FK =R,, +R,
ZM” =0=R, *ETSL-FK *LK

E ISL

K

 

geomeny=AT=AH+(AK—AH)*

ETSS=KT=£= FH+FK_RH 3 FK*LK2
AH+(AK_AH).§ZS_£ I...2 *(AK—AH)+LK *LT‘AH

K

 

 

182

REFERENCES

. Hubbard, R., and C. Gedraitis, “Initial Measurements and Interpretations of Seat
Factors for the ASPECT Program,” Society of Automotive Engineers (SAE) Paper
Nmnber 1999-01-0958, presented at the 1999 SAE International Congress, Marcln,

1999.
. Working Model® 2D Software. Knowledge Revolution. Version 4.0 ©1992-1996

. Excel Software. Microsoft Corporation. Ofice 2000 Version

. “Devices for Use in Defining and Measuring Vehicle Seating Accommodation”,
Society of Automotive Engineers Recommended Practice 1826.

. Schneider, L., M. Reed, R. Roe, M. Manary, C. Flannagan, R. Hubbard, and G. Rupp,
“ASPECT: The Next Generation H-point Machine and Related Seat Desigrn and
Measurement Tools”, Soc. of Auto. Engin. (SAE) Paper Number 1999-01-0962,
presented at the 1999 SAE International Congress, March, 1999.

. Hubbard, R., and N. Radclifi‘c, “Seat Pan Load-Deflection Measurements For
Determination of Seat Factors with the ASPECT Butt-Thigh (ABT) Segment of the
ASPECT Manflcin Prototype at MSU and the SAE 1826 Manikin,” submitted to
TecMath Inc., reprinted in Appendix B.

. Reid-Bush, T. “Posture and Force Measures of Mid-Sized Men in Seated Postures,”

dissertation for the degree of PhD. Dept. of Materials Science and Mechanics
Michigan State University, 2000.

183

   

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