TOB cooling pipework

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TOB cooling pipework
TOB note 05/1
8/13/2017
TOB cooling pipework
A study of the mechanical behaviour of a solder joint between a
brass connector and a copper pipe
Christian Lönnqvist
Helsinki Institute of Physics
Abstract
The strength of a soft tin-led-silver soldered joint between a brass connector and a copper
pipe was tested in a tension test bench. For this purpose a brass connector end, equal to
the ones’ of the connectors used on the CMS Tracker Outer Barrel (TOB) to connect the
cooling liquid feed pipes to the cooling pipework, was designed and manufactured. The
copper pipes used, had a wall thickness of 1 mm. Analytical calculations were made for
estimating the specimens’ fracture force, as well as for understanding the stress
distribution in the solder joint. The two test specimens were loaded to fracture in a
standard tension test bench, equipped with a 20 kN load cell. The intermediate fracture
force was very large in magnitude, approximately 4 kN, in the middle region of were the
two different analytical mathematical calculation models for shear stress in the solder
predicted it to be. Microscopic observations were made, which proved that the analytical
mathematical calculation model predicted the fracture to start where it did. Fracture is
always in the solder material.
Table of Contents
1.
2.
3.
Introduction ................................................................................................................. 3
Design and manufacturing of test pieces .................................................................... 5
Material mechanics ................................................................................................... 11
3.1
Conclusion on the analytical calculations ......................................................... 11
4. Test procedure and results ........................................................................................ 13
5. Discussion ................................................................................................................. 17
6. Conclusions ............................................................................................................... 21
References ......................................................................................................................... 22
Appendix 1 ........................................................................................................................ 23
Appendix 2 ........................................................................................................................ 24
Appendix 3 ........................................................................................................................ 25
Appendix 4 ........................................................................................................................ 26
Appendix 5 ........................................................................................................................ 29
Appendix 6 ........................................................................................................................ 31
Appendix 7 ........................................................................................................................ 33
Appendix 8 ........................................................................................................................ 38
Appendix 9 ........................................................................................................................ 43
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1. Introduction
The CMS tracker has plenty of cooling pipes. These pipes are needed for the
transportation of coolant to detectors and cables dissipating some 40 kW of heat in the
CMS Tracker Outer Barrel (TOB), Tracker Inner Barrel/Tracker Inner Disks (TIB/TIDs)
and Tracker End Cap (TEC).
The main components of the TOB cooling pipework are supply pipes, connectors,
manifolds and rod cooling pipes. The cooling pipework shall be manufactured in
materials suitable for the tracker's operating environment and for the load cases that will
act on the pipework. The sensitive silicon detectors, as well as the clean room
environment, put their own restrictions on what manufacturing methods can be used for
the assembly of the cooling pipework.
The chosen materials for the TOB cooling pipes and manifolds are copper-nickel (70%
Cu, 30 % Ni) pipes, and brass connectors. The chosen material for the TOB cooling feed
pipes is soft copper (Cu DHP-O). The joining is made using brazing and soldering. When
joining pre-assemblies of Cu-Ni manifolds and pipes, and brass connectors, brazing is the
most reliable joining method. Brazing is, however, not applicable in the in-situ joints
needed to be done at certain parts of the cooling pipework. This is because the brazing
method requires process temperatures above 450 °C needed to heat up the base and filler
materials. This temperature is usually achieved with a gas burning flame, which can not
be used in the near vicinity of the silicon trackers. Another reason for why one can not
use brazing as an in-situ joining method, is the extensive cleaning process one should
apply to the pieces for removing the aggressive residual brazing flux that is used to
remove oxide skins from the assembly pieces on the joining surfaces. [1]
The in-situ assembly of the cooling pipework includes two types of pipes, the cooling
connection Cu-Ni pipe (from circumferential cooling manifold pipe to each rod cooling
pipe) and the copper feed pipe (from cooling feed-pipe at the TOB end flange to the
circumferential cooling manifold pipe). The best found method for manufacturing leak
proof joints of these connections, is to apply soldering with solder with mild, solid
multicore flux. The advantage of soldering is the low temperature needed for melting the
solder filler, typically around 200 °C (183 °C for eutectic tin-lead solder) and the
relatively strong joint that can be achieved with this joining technique. An advantage is
also the use of a more mild flux, requiring less tedious cleaning methods for removing it.
The low melting temperature also allows the use of an electrical solder iron, which is of
advantage when working in a sensitive environment.
The soldering joint between Cu-Ni pipes and brass connectors has been studied
extensively at CERN/EP/TA1 [2]. However, there has not yet been a clear understanding
of the feasibility and mechanical behavior (i.e. joint strength) of a rather large (Ø10 mm)
solder joint, between the copper feed pipe and the brass connector. For this purpose a Cupipe - brass connector test setup was designed. This specimen was then tested in a
standard tension test apparatus, from where load - displacement curves could be
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retrieved. Before testing, some basic analytical calculations were made for the evaluation
of the strength of the soldered joint.
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2. Design and manufacturing of test pieces
The design of the setup solder joint was to be as identical as possible with the design of
the actual joints of the pipes and connectors for the TOB (as shown in appendix 1.). For
this purpose dimensions for the end-part (nr. 1, figure 1.) of the brass connector were
directly taken from the drawing of the final connector. This end-part was then fit to a
clamping surface (nr. 2, figure 1.) suitable for the attachment in the tension test apparatus.
To avoid the heat to flow away from the soldering region, the clamping surface with its
large mass concentration was displaced far from it. The brass1 connector was made of the
same material as for the corresponding brass connectors for the TOB, available as a
round bar. A drawing of the brass connector for the setup can be found in appendix 2.
Normal machining procedures like turning, milling and drilling are sufficient for
achieving the final geometry of the brass connector.
1
.
2
.
Figure 1. Brass connectors awaiting the last machining of the small vent holes.
The joint between the copper pipe and brass connector is 8.4 mm long and has 0.05 mm
radial clearance into which the molten solder is driven by capillary forces (see figure 2.).
The rest of the test setup connector end is designed to the wall thickness of 0.8 mm to
match the corresponding connectors used in TOB. To ensure escape of solder gases
during soldering vent holes consisting of a 3 mm hole in the axial direction and a 1.5 mm
hole in the radial direction, had to be made (see figure 3.). This additional 3 mm hole is
needed, because drilling the 1.5 mm hole into the 0.8 mm thick connector wall, could
weaken it and affect the test results.
1
Brass ISO 1637-HB (CuZn39Pb3 Hard), drawn round D 25 mm. CERN SCEM number: 44.09.07.125.8
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Figure 2. Location for where the solder filler material is added during soldering
Figure 3. A vent hole to free soldering gases from inside the connector during soldering
The whole setup consists of two brass connectors and one 70 mm long, Ø10 mm and 0.5
mm thick soft copper pipe2. Important to notice though, is that the two first setups were
tested with soft copper pipes with a wall thickness of 1 mm. A common tin-lead-silver3
alloy was used as solder filler in the joints. One of the two first setups was soldered with
flame, and the other with an electrical soldering iron. A drawing on the setup can be
found in appendix 3. A picture of the setup in the tension test apparature can be found in
figure 5.
Before soldering, the pieces were carefully cleaned with alcohol, as seen below in figures
4. and 5.
2
3
Soft copper pipe D 10 ISO 1635, s 0.5. CERN SCEM number: 39.71.05 B
Sn62Pb36Ag2. CERN SCEM number: 29.20.01.349.6
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Figure 4. Cleaning the brass connector with alcohol
Figure 5. Rubbing the oxide layer of the copper pipe with an emery cloth
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Figure 6. Cleaning the copper pipe with alcohol
Figure 7. Parts ready for assembly
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Figure 8. Aligning the setup
The setup was aligned and then temporarily glued on a pair of plastic bars (see arrow on
above figure) to assure clearance and isolation from the table.
Figure 9. Heating up the joint region with a soldering iron which has a half-circle
shaped head to best match the round connection piece
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The joint region needed to be pre-heated, to assure that the joint would be as wetted as
possible with solder filler, and to decrease the risk of a cold bond. For this a custom
shaped soldering iron was used to best match the joined pieces.
Figure 10. Adding solder filler
As filler, a solder with mild, solid multicore flux was used.
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3. Material mechanics
In order to get an estimation on the magnitude of the fracture force of the brass connector
and copper pipe test setup, before being pulled apart in the tension test, some basic
calculations were made. These calculations can be found in appendices 4,5,6,7,8 and 9.
They evaluate the fracture force in the copper pipe, in the brass connector and in the
solder joint. For the evaluation of the fracture force in the solder joint there has been used
two calculation methods, one assuming constant distribution of the shear stress over the
joint, and another assuming exponential distribution of the shear stress over the joint. The
calculation giving the highest stress peak (either in one of the cross sections or in the
solder joint) will be the one determining the fracture force of the setup.
In the below table one can find values for different material constants used in the
calculations. These values have been gathered from various sources [3,4,5], from CERN
material stores catalogue or from the internet. Careful criticism has been applied when
choosing them.
Table 1. Material constants
Young's modulus, E
(GPa)
Shear modulus, G (GPa)
Poisson's ratio, ν
Ultimate
tensile
strength, σts (Pa)
Yield strength, σys (Pa)
Shear strength, τss (Pa)
Brass
(CuZn39Pb3)
Copper pipe Solder
(soft
(Sn62Pb36
annealed)
Ag2)
97
115
33
37
0.31
44
0.31
12.7
0.4
415
345
48.3
140
240
115
195
52
3.1 Results of the analytical calculations
Different stress components will be dominative in different parts of the setup when
loaded. Pure tension stress will dominate in the cross sections of the Cu-pipe and the
brass connector. Shear applied by the relative difference in axial elongation between the
Cu-pipe and brass connector will be dominative in the solder joint. This solution is
derived from a model commonly used for glue joints [6]. Therefore some criticism should
be applied towards the results achieved with it. The simple calculation, assuming constant
shear over the joint, implies no elongation in the Cu-pipe and the brass connector under
loading. Failure forces in different parts of the setup are presented below.
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Table 2. Calculated failure forces
Cu-pipe cross sect. (N)
Brass conn. cross sect. (N)
Joint, constant shear (N)
Joint, exp. shear (N)
0.5 mm Cupipe thick.
5149
9179
13068
1080
1 mm Cupipe thick.
9755
9179
13068
1018
The setup with 0.5 mm thick copper tube is expected to fail at either Cu-pipe cross
section or at the solder joint. The exponential shear stress calculation yields a lower
fracture force than that for the Cu-pipe cross section, however, the glue model used in the
shear calculation is not fully applicable for solder joints, because it does not account for
plastic stress relaxation that can occur in metals. Therefore, one can not be confident
about in which region between 1080 N and 13068 N the fracture force really will be. The
setup with 1 mm thick copper tube is expected to behave in the similar way. Interesting to
notice though, is that the calculation assuming exponential shear stress distribution
predicts a failure force with one order of magnitude less than the one that assumes
constant distribution of the shear stress. This is because the solder joint will get the
highest distortion close to the ends, where the relative elongation of the copper tube and
brass connector are at their highest in magnitude over the joint, hence implying high local
shear stresses at these points. See figure 4. below.
Figure 11. Calculated shear stress [Pa] in the solder joint using the mathematical model
of a glue joint, x [m] runs in axial direction.
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4. Test procedure and results
The tension test apparature had to be equipped with the proper load cell and fastening
clamps. For this setup the tension test bench was equipped with a 20 kN load cell. Below
one can see a picture of the test bench.
Figure 5. Tension test bench
A pair of distortion measurement hooks were also installed, see figures 6. and 7. These
are used when one wants to measure the distortion from one exact portion of the tested
object. The use of these hooks proved to be unsuccessful though, because they did not
clamp correctly on to the surface of the test object.
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Figure 6. Distortion measurement hooks
Finally the test piece was erected into the tension test bench. Even though careful
adjustments were made for aligning the piece with vertical axis of the machine, one could
expect that some bending stresses were transferred to the test piece.
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Figure 7. The test specimen in the tension test bench
Because both test pieces were non-standard, they were loaded at a low 1 mm/min speed.
When loading test piece 1. (soldered with electrical heating iron), only the fracture force
was retrieved from the test, because of unsuccessful use of distortion measuring hooks.
When loading test piece 2. (soldered with flame) the tension test machine's own
distortion measuring sensor was used and proper load-displacement curves were
retrieved, as well as the fracture force. The test results as well as the load-displacement
curves are presented below.
Table 3. Test results
Loading speed (mm/min)
Fracture force, (N)
Fracture displacement, (mm)
Test piece 1.
(soldering iron)
1
3784
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Test piece
(flame)
1
4342
4.963
2.
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Figure 8. Load-displacement curve of the test piece 1. The curve shows the slipping of the
distortion measurement hooks.
Figure 9. Load-displacement curve of the test piece 2.
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5. Discussion
Both test setups failed at the solder joint at about 4 kN. This might or might not be a
surprise. The first thing one can conclude is that, the magnitude of the fracture force for
the solder joint is in the middle region between what the two analytical calculation
models describe it to be. Thus in reality the stress distribution is not constant, like the
simple model describes it, on another hand, the distribution seems more even than how
the glue-joint model describes it to be. Some error, compared to the test results, will also
be added from the fact that additional shear is added from bending (due to missalignment). The mathematical model of the joint is also erroneous because it is treated
mathematically as a flat plate, while in reality it is a tube. This means that in a highly
curved joint, as in this case of a small diameter tube, part of the shear is not counted for
in the analytical mathematical model.
As one can see from the curve, the setup's load-displacement behavior is essentially
linear up to the loading of 1,5 kN. The displacement at this point is a superposition of the
displacement in the brass connector and in the copper tube. One can neglect the axial
displacement in the solder joint. So we can write (index 1. refers to brass and index 2.
refers to copper):
 appl.1 
F
A1
 appl.2 
F
A2
 mat.1  E1 1
 mat.2  E 2  2
And if we equalize the applied stress with the stress acting in the material, and put
F=1500 N, we get:
1 
F
 0.7  10 3
E1 A1
2 
F
 0.46  10 3
E 2 A2
 tot   1   2  1.15  10 3
When we have the elongation, we can calculate the stresses in the materials:
N
 0.7  10 3  67.9  10 6 Pa
2
mm
N
 2  E 2   2  115  10 9
 0.46  10 3  52.9  10 6 Pa
2
mm
 1  E1   1  97  10 9
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As we can see from the load-displacement curve, the test setup starts yielding at a loading
of 1500 N, then this yielding has to be accounted for from either the copper pipe or the
brass connector. The above result shows that the magnitude of tension stress in the brass
is higher than in the copper, but it is still far from the yield stress for this alloy. The
tension stress acting in the copper, is close to the half of its yield stress though. Based on
this, and on the fact that copper has a vague definition of its yield strength which is very
strongly dependent on the exact temper of the material, one can conclude that the copper
pipe is the one to yield first. The next question to ask is, will the brass alloy yield in the
cross-section at all? This we can find out when we know the maximum load (4342 N)
applied on the setup in the test.
 1 max 
Fmax
 2.02  10 3
E1 A1
 1 max  E1   1 max  196.3  10 6 Pa
From this, we can conclude that the tension stress in the brass connector will not reach
the yield stress of the alloy during the tension test. Thereby the deformation, and all the
yielding, is happening in the copper pipe section. The shape of the load-displacement
curve is essentially dominated by the material behavior of copper. A clear effect of work
hardening can also be seen from the load-displacement curve. This is typical material
behavior for copper when getting cold worked.
Upon the above reasoning one can understand why the samples fractured in the solder
joint. During loading, the large magnitude of elongation in the copper pipe and the
smaller magnitude of additional elongation in the brass connector distorted the solder
joint in axial direction by the magnitude of the relative difference in the elongations. This
implied, in magnitude, large shear stresses in the solder at both ends of the solder joint.
This is what the model for distributed shear stress, more qualitatively than quantitatively,
already predicted. This fact can also be verified from microscopic observations (see
pictures below).
Figure 10. Locations from where the microscopic pictures were taken
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Figure 11. Fracture in solder joint
Figure 12. Fracture in solder joint
Figure 13. Fracture in solder joint
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Figure 14. Fracture in solder joint
These are pictures of the solder joint in the fracture ends of the two test samples (see
figure 10.). One can clearly see that the fracture has initiated at the connector end of the
solder joint (where the photos are taken of), where the magnitude of the elongation of the
copper pipe has been the largest. The fracture is always in the solder material. No fracture
can be found in the material interfaces or in the copper or brass walls.
From the test results, one can see a small difference in the magnitude of the fracture force
(558 N) of the two samples. It is difficult to draw any definite conclusions on what causes
the difference, because the test series consisted of only two samples. The difference
might evolve from different amount of solder used (the joint of sample 1. was not entirely
wetted by the solder) as well as from temperature differences during the soldering
(sample 1. was soldered with electrical soldering iron, while sample 2. was soldered with
a flame). It can also evolve from differences in alignment of the pieces in the setup,
transferring different magnitudes of bending stresses to the solder joint.
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6. Conclusions
One can conclude that this type of solder joint is very resistant to loading indeed. Forces
in magnitudes of what is required to break the solder joint between the brass connector
and the copper feed pipe, about 4 kN, will most likely not act under normal operation
situations of the TOB, and neither during the assembly of the pipe work. Thermal
expansion and shrinkage due to temperature differences between load cases during
operation of the TOB will induce loads on the pipework to some extent. The magnitudes
of these forces are essentially dependent on how flexible the pipework is. Even though
these loads might not be severe to the solder joints, one should bare in mind, that they
might cause damage somewhere else. Therefore one should strive to decouple degrees of
freedoms by designing and manufacturing an appropriate pipe supporting.
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References
1. Based on discussions made with Antti Onnela (CERN/PH/DT2) during summer
2005.
2. Antti Onnela et al., Soldered low-mass cooling pipe connections for LHC
trackers, March 2001
3. http://www.efunda.com/materials/solders/tin_lead.cfm,
Tin-lead
solders,
8/26/2005
4. http://www.matweb.com/, Material property data, 8/26/2005
5. Reijo Kouhia et. all, Constitutive models for stress analysis of solder joints,
http://www.tkk.fi/Current/NSCM12/proceedings/kouhia.pdf, 8/26/2005
6. Einar Halmøy, Kompendium i liming, fall 2003
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Appendix 1
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Appendix 2
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Appendix 3
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Appendix 4
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Appendix 5
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Appendix 6
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Appendix 7
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Appendix 8
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Appendix 9
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