UNIVERSITY OFCALIFORNIA

Assessment of local compliance in joints and mounts of global support structures :Testing and Analysis of the STAR Inner Detector Support

June 19, 2013Neal Hartman on behalf ofJoseph Silber, Eric Anderssen and the LBL Team

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Overview

Brief description of the HFT upgrade at STAR, and the new inner global support structure, the IDS (“Inner Detector Support”)

Early deflection analysis results (overly optimistic)

Testing of the IDS for model verification (a scheduled milestone)

Re-analysis (detailed discussion)

Lessons learned

Time permitting - Brief discussion of HFT’s pixel insertion system, and implications for complexity / mass

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STAR HFTHFT is a new inner tracking system for

STAR, with 4 layers of silicon and 6 gem disks

Timeline:• Nov 2011 main support structures +

FGT were installed• July-Dec 2012 PXL support and

PXL for engineering run• Summer 2013 full PXL + IST + SSD

Key component is PXL:• 2 innermost silicon layers• Truly rapid insertion/removal• Very low mass• TPC is great; PXL will much

improve pointing

At LBL we’re building / have built:• All the support structure (IDS)• All of PXL• IST local supports

TPC Volume

OFCOuter Field Cage

IFCInner Field Cage

SSDISTPXL

HFT

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STAR HFT Inner Detector Support (IDS)

5

4.6 m

0.8 m

Structure Mass = 35kgApplied Load = 200kg

WSC/ESC Mandrel WSC/ESC Layup

Cone Layup Flange Layup

Insertion Rail Bonding Assembled Structure at LBNL Just before insertion at BNL

Cone Machining Flange Bonding

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Initial FEA model

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Careful accounting of masses

Modulus properties based on multiple tensile tests for the composite materials

Layered anisotropic composite properties fully specified, ply-by-ply

Quasi-kinematic BCs at designed locations

Face-face contacts at joints

Apparently good nominal performance:607 μm max vert defl @ full load)

And yet…

Pre-test predictions Measured deflections

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-6.000

-5.000

-4.000

-3.000

-2.000

-1.000

0.000

1.000

2.000

57.000 117.395 159.405 218.820 249.075

Defle

ction

(mm

)

Applied Central Load (kg)

2011-09-23 IDS Test Results

OSC +Y

OSC -Y

OSC +X

OSC -X

WSC +Y

ESC +Y

WSC -X

ESC -X

WSC Ztop1.88 mm @ 250kg central load

5.23 mm @ 250kg central load

Predicted deflection only 36% of actual measured value in the “model verification” test

(Note: the 250kg central load applied during test is equivalent to > 2x the distributed 194kg installed load.)

Load Test Setup(Model Verification)

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4.63mWest Support Cylinder (WSC) East Support Cylinder (ESC)

Outer Support Cylinder (OSC)

West Main Flange(bonded-in aluminum ring)

East Main Flange(bonded-in aluminum ring)

East ConeWest Cone “Saddles”

Weight Support Plate

Support table

End SupportsEnd Supports

CENTRAL LOADINGTHRU SADDLE

Kinematic turnbuckle supports(same as used in final install)

Horizontal (1x)(both sides)

Vertical (2x)(both sides)

Axial (1x)(one side)

Test does not use actual, distributed, installed loads.Test has a significantly higher load, centrally located.

Many “DOF” measured during testDial test indicators were placed at 18

locations about the structure during test – avoiding prejudice as to which would be redundant, which important

Thus capturing a rich description of how the real object deforms

This data was invaluable in diagnosing the model, particularly because

• the angled flat feature of the ESC and WSC causes sideways coupling with the vertical deformation

• the mounting points are quasi-kinematic, due to constraints on assembly procedure

• multiple bolted joints, where bolt preload is necessarily non-optimal, due to material constraints

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angled flat

Diagnosis of the model

Identified as many possible sources of compliance and independently studied each; then integrated all into the global model

• Must identify sources bottom-up and close loops with independent tests, not twiddle with parameters at the top

A perfect model would have an additional amount of compliance over the original equal to

measured/orig.model – 1 = 5.258/1.881 – 1 = 1.79Immediately identified one simple BC error, accounting for 23% of

the total error (0.41 out of the 1.79)

The rest of the compliance sources are more subtle and interesting

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Integration of the independent compliance sources

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3.1.1 Double boundary condition in Z

In the design, the axial (Z) direction is constrained at one point only

In the model, an extra erroneous constraint had been applied on the opposite side of the structure

Removal of extra constraint accounts for 0.41 of 1.79

• (again, this is an arbitrary unit indicating normalized missing compliance)

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Actual Z=0 constraint here

Erroneous Z=0 constraint here

3.1.2 Overconstraint boundary condition in Y

A more subtle and interesting BC error

Design requirement that any one vertical support turnbuckle be removable at any time• Intentional state of overconstraint, for assembly purposes

In reality, when 4 vertical supports are present:• only 2 (at diagonal corners) carry the load (because torsional stiffness of structure is so high)• 3rd support stabilizes structure with essentially no load• 4th is free to deflect to any vertical position (within constraints of thread pitch clearance, etc)

Examination of deflection data for turnbuckle supports (wisely recorded during load test) showed that NW and SE supports carried the weight load, with NE the stabilizer

Removal of SW vertical support from model 0.12 of the 1.79 compliance error14

Horizontal (1x)(both sides)

Vertical (2x)(both sides)

Axial (1x)(one side)

3.1.3 Finite stiffness of vertical constraints in test fixture

In original model, vertical supports treated as perfect point constraints in the Y direction

Calculating the stiffness of these turnbuckles a priori would be complex, due to thread-on-thread and spherical bearing interfaces

From turnbuckle deflection data, however, could extract real, accurate stiffness of 3e6 N/m per support -- 0.406mm @ 250kg , directly contributing 0.07 of 1.79 to global compliance

With good stiffness values, simple to replace vertical constraints with spring elements in the model (and model then confirmed that only the diagonals actually carried load)

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Trial #1, 2011-09-21Location Description 0 57.216 117.631 159.641 219.081 249.336WSC -Y2 Bottom of west main flange 0 0.4318 0.9398 1.4478 2.2606 2.3622WSC -Y3 Northwest turnbuckle rod-end 0 0.0254 -0.0254 0.0254 -0.3302 -0.4064ESC -Y2 Bottom of east main flange 0 0.3048 0.6858 0.9398 1.3208 1.5748ESC -Y3 Northeast turnbuckle rod-end 0 0 0 0 0 0ESC +Y2 Top of ESC shell near main flange 0 -0.0508 -0.1524 -0.2159 -0.3302 -0.4191ESC -X2 Middle horizontally and vertically of ESC shell 0 -0.0254 -0.0508 -0.0635 -0.0889 -0.1143

3.01E+06 N/m <-- spring stiffness for 250 kg force / 2x WSC -Y3 deflection

Load (kg)

<-- shows that a turnbuckle wasn't loaded

<-- loaded turnbuckle deflection

Deflection (mm

)

3.2 Bolted Joint

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Key CFRP bolted joints

Strain plot indicates how joint/flange locations are crucial to global deflection

Revised model joint geometry,bolts and thick sections all solid modeled, parameterized stiffness of contact faces(more discussion to come)

CONE

OSCOriginal model joint geometry,face-face contacts at joints,some thick sections modeled as shells

OSC

CONE(not shown)

Note:Simply replacing face-face contacts with point contacts at hole locations would increase compliance 9% (separate study). Still insufficient to capture compliance accurately.

Nearly equivalently, can model as 6 DOF springs

Picture shows original FE model stackup of CN60 plies in the cone

Original model treats entire cone as shell body in this way

But as approach edges, we have a thick section deforming in modes where typical thin wall, in-plane assumptions do not apply

The middle section, even though 4mm thick, is ok to model as shell

At edges, where discrete joint loads effective, the shell assumption breaks down

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3.2, Bolted joint grip integral with cone

FLEXUREFL

EX

UR

E +

TO

RS

ION

+ T

HR

U-

THIC

KN

ES

S

NO

RM

AL

LOA

DS

FLEXURE + TORSION + THRU-

THICKNESS NORMAL LOADS

3.2.1 Bolted Joint Nonlinear Contact ModelBolted joints thru CFRP grips: key point of compliance in the overall structure.

To quantify this compliance, local model of the joint analyzed with high mesh refinement and nonlinear contact definitions.

“Nonlinear” evaluate over a series of substeps to capture the changing contact stiffness as the grip rotates and compresses.

Computationally very expensive, and adverse on convergence. Global model still needs to be linear, so make a linear approximation of the joint based on the nonlinear local model (next slide).

The effect of changing from a standard, high normal stiffness bonded contact to this nonlinear model is 39% increase in compliance (0.39 of 1.79).

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OSC ½ shell

Refined grip model

Bolt preloading

Central loading

3.2.1 Linear approximation of nonlinear joint model

Via contact stiffness parameter, method essentially includes a spring in the flange – the separate nonlinear model sets the stiffness for the larger linear model

Abscissa at right shows varying the normal stiffness factor of a linear joint model.

Ordinate is correlation to the nonlinear model

Resulting contact stiffness parameter: 0.002, with a sensitivity of 12% for every Δ of 0.001

This parameter is the linear “knob” one can bring into the global model, to approximate the nonlinear effects.

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• Disadvantages of this approach:• Decouples joint design from global model• Obscures joint behavior in global model

• Advantages of this approach• linear and convergent• far more accurate than a bonded contact

3.2.2 Wedged joint gap

Partial wedge-shaped gap observed at the OSC bolted joint during assembly of the IDS

Caused by a springback effect after facing the contact surfaces under clamped conditions in the mill.

Gap up to approximately 0.2mm in some locations.

To quantify effect on global joint stiffness, a nonlinear contact joint model run with a partial wedged gap.

Plot on right shows joint closure (nonlinear model) under bolt preload

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15°

R 200mm

10mm

6mm

Wedge gap0 to 0.2mm

2mm flat (no gap)

3.2.2 Wedged joint gap

Top plot:measured bolt torques (confirming preload repeatability) for titanium M4 threaded fasteners in IDS

Bottom plot:nonlinear local model’s predictions of gap’s effect on global stiffness

• Depends on stiffness of adjacent bonded member

• For the situation at hand, it’s 7-10%

• i.e., 0.07 of the 1.79 compliance error

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3.2.3 Grip material anisotropy

Grip of the bolted joints modeled as solid body, due to low aspect ratio

Approximated as isotropic in original model, since layup is quasi-isotropic

In reality, anisotropic (transverse isotropy), with much reduced moduli thru the thickness in normal (EZ) and shear (GYZ, GXZ)

Laminate shear moduli bracketed on the high and low end by a unidirectional ply’s out-of-plane shear moduli:

• 2.1 GPa = G23 ≤ GYZ = GXZ ≤ 6.3 GPa• So value of 4 GPa chosen as reasonable

Conversion to transverse isotropy 22% increase in global compliance (0.22 of 1.79)

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3.3.1 CPT of woven laminate

Generally for thin laminates, CPT isn’t important for mechanical FEA – just needs to agree with the moduli inputs

IDS however, has thick CN60 weave laminates at key locations (cone, bolt flanges)

And in these thick sections, flexure is dominant deflection mode

Original model had CPT = 264 μm, calculated a priori before test objects existed

Subsequent measurements showed CPT to be 250 μm.

This 5.3% lesser thickness indicates flexural stiffness reduction of 1 - (94.7%)³ = 15%(0.15 of 1.79)

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Thickness measurements of as-built parts

3.3.2 Woven in-plane stiffness reduction factor

FEA model treats each CN60 woven ply as stacked pair of half-thickness uni plies at right angles

In original model, out-of-plane undulation of the weave not accounted for. In reality, it’s significant.

• Typical engineering rules of thumb call for reduction factor of 5% to 15%

3pt and 4pt tests made of CN60 test beam to assess

Also 6061 reference beam on same setup, to confirm test validity

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Tests gave woven stiffness reduction factor = 13%(0.13 of 1.79)

Note in table below that an FEA of the beam using original model definitions versus corrected properties is 30% stiffer.

This agrees well with simply combining the 15% CPT reduction with the 13% factor due to weave

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Woven in-plane stiffness reduction factor

mass (g) 179.26 mass (g) 167.87 mass (g) 245.18 mass (g) 179.26 mass (g) 167.87mass err 7.1% mass err 0.3% mass err 0.4% mass err 7.1% mass err 0.3%

applied mass

applied load deflection

zeroed deflection deflection

zeroed deflection deflection

zeroed deflection

applied mass

applied load deflection

zeroed deflection deflection

zeroed deflection

kg N mm mm mm mm mm mm kg N mm mm mm mm0.0 0.00 0.040 0.000 0.048 0.000 0.145 0.000 0.0 0.00 0.040 0.000 0.048 0.0000.1 0.98 0.075 0.036 0.094 0.046 0.239 0.094 0.1 0.98 0.071 0.031 0.089 0.0410.5 4.91 0.218 0.178 0.278 0.230 0.617 0.472 0.5 4.91 0.197 0.157 0.251 0.2031.0 9.81 0.396 0.356 0.508 0.460 1.089 0.944 1.0 9.81 0.354 0.314 0.454 0.406

1.181 11.59 0.410 0.371 0.527 0.479least-squares stiffness (N/mm) --> 27.567 21.331 10.388 31.263 24.186

R ² --> 1.00000 1.00000 1.00000 1.00000 1.00000

FEA variation from measured value --> 34% 4% 2% 35% 5%

3-point bend 4-point bendCorrected Defn6061-T6 Original DefnOriginal Defn Corrected Defn

3.4.1, 3.4.2 Minor effects

Changed bonded contact joints throughout model from “pure penalty” contact formulation to slightly more computationally expensive Lagrangian method

• increases global deflection 3% (0.03 of 1.79)

Original model didn’t include 2x ø100mm access holes in each cone• increases global deflection 1% (0.01 of 1.79)

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Integration of the independent compliance sources

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Results after incorporating all changes into the model

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MEASURED CENTRAL DEFL

ORIGINAL MODEL FINAL MODEL

Also note that the other deflection terms (besides central) agree much better in the final model.(Exception “ESC –X”: after much study this is believed to have been an anomalous bottoming-out of indicator probe tip during test – not included in weighted average figure of merit)

Summary / lessons learned…There is a big difference between nominal initial studies and eventual as-built performance:• Original FEA with installed loads: 0.607 mm central deflection• Corrected model (with benefit of physical test): 2.501 mm

Careful deflection test of final global support essential; plentiful distribution of indicators very useful• Redundant measurements not only error check• They also identify rigid body modes vs deformations

As early as possible, key BC support points in model should be modeled as spring elements with physically-tested stiffnesses

Take nothing for granted in modeling of bolted joints with CFRP grip – the low transverse stiffness of CFRP + low allowable preloads can defy our usual engineering intuition

Subtle CPT effects become important to quantify for thicker cross-sections in bending• In a larger sense, the D-matrix becomes important

In-plane stiffness reduction factors for weave can be significant, even with excellent compaction / process control

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Takeaway point:Most of the simple geometries, tubes and plates and so forth, are almost rigid bodies in our application. The materials are that good. But it is the difficult-to-model components – interfaces, bolted joints, low aspect ratio flanges – which cause a large amount of the deflection.

Clam-shelled Insertion Mechanism

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HFT pixel insertion mechanism

Rapid (< 1 working day) replacement of innermost 2 pixel layers

Clamshelling of 2 halves in around pipe as approach IP – hinged, tracked design

Cantilevered from kinematic mounts (engaged at end of insertion)

Air cooling makes service routing much easier

Somewhat similar radii (26mm, 82mm) as Cartigny

Trade-off is complexity and significant mass at Z ≥ 575 mm on east side

Rough scaling guess as to mass if one were to implement something like this at ATLAS size:

• 870 g of carbon fiber @ Z ~ 800 mm• 3640 g of aluminum @ Z ~ 1100 mm• Might be able to swap some CF for alum

~2250 g CF, ~1450 g alum

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STAR: PXL Insertion Test-bed

Hinge mechanism

Dovetail plate

supports pixel staves

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STAR PXL Kinematic MountsDesign constraints

• 50μm positioning repeatability for all parts (6x detector halves)

• Insertion is remote (detector tracks in ~3m along circuitous route before engaging mount); no tool access

• Confined space, low mass, nothing magnetic

• High enough retention force to keep detector stably located, but low enough to insert/remove detector without significant impact

One could design a mechanism with rotating components, i.e., cams and locks; chose the other route, to make it a flexures-with-friction problem. This allows:

• Simple FBD analysis (as long as you test μ first!)

• Accurate spring stiffnesses machined-in to parts by design

Upper kinematic mount (XY or XYZ) Lower kinematic mount (X)

PXL Support Half engaged in master tool

Test Stand

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STAR PXL Kinematic MountsForce-Disp Results on Test Stand

-30

-20

-10

0

10

20

30

0 2 4 6 8 10 12 14 16

Inse

rtion

/Ret

racti

on Fo

rce

(N)

Position (mm)

0.0 0.5 1.0

1.5 2.0 2.5

FBD predicted max insertion force FBD predicted max retraction force

Applied weights (kg)

μ = 0.2

Insertion

Retraction

EN

D S

TOP

“click-in”

“break-out”

Calculated max (μ = 0.2)

Calculated min (μ = 0.2)

TOP EAST KIN MOUNT, TEST 2011-08-29

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Rough guess as to how HFTs pixel insertion mechanism’s mass would scale to something of ATLAS size

Insertion MechanismVery rough estimate, IF an insertion mechanism were to be used.All masses taken from STAR HFT's PXL insertion mechanism.

2.7 g/cm³, aluminum density1.7 g/cm³, carbon fiber density

1300 mm, length of I-beam710 mm, length of PXL (carbon fiber structures)

183% size scaling, by length180 g, I-beam total mass, approx

87 g, PXL total mass, approx207% size scaling, by mass195% average size scaling

Assume aluminum if not CFPart Volume STAR, CF? ATLAS, CF? Carbon Fiber Aluminum Carbon Fiber Aluminum

- cm³ 0 or 1 0 or 1 g g g gD_terminator_extend_to_hinge_v2 85.2 0 1 0.0 230.1 144.9 0.0hinge_leaf_in 62.4 0 1 0.0 168.4 106.0 0.0hinge_leaf_out_blistered_v2 60.7 0 1 0.0 163.9 103.2 0.0release_hinge 84.9 0 0 0.0 229.1 0.0 229.1hinge_D_end_v2 33.3 0 0 0.0 90.0 0.0 90.0transition_tube_cyl_v4_u 130.6 1 1 222.0 0.0 222.0 0.0kin mounts (3x) 11.6 0 0 0.0 31.3 0.0 31.3pins (misc) 2.3 0 0 0.0 6.3 0.0 6.3bolts (misc) 5.7 0 0 0.0 15.3 0.0 15.3

Total: 222.0 934.4 576.1 372.02x total (both halves): 444.1 1868.8 1152.3 744.0

2x, with size scaling: 865.9 3644.2 2246.9 1450.9

STAR Mass ATLAS Mass

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Backup

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IDS Loads and BCs, corrected arrangement

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Method of summing the compliances in the integration of separate effects

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Distributed loading of IDS (installed loads), and oversizing of central test loads

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+ZWEST EAST

48.3kg@ Z = +870mm

145.8kg@ Z = -1446mm

Fc/2@ Z = +556.3mm

Fw@ Z = +2316mm

Fe@ Z = -2316mm

Fc/2@ Z = -556.3mm

EQUIV. CENTRAL LOADEQUIV. WEST LOAD EQUIV. EAST LOAD

(installed) (installed)

• Load test was performed with 250kg central load

• In fact, if solve the force diagram above, the correct test load for 150% proof should have been• 180kg central load, plus• 110kg end load on east side only

(to get the moments correct)