IPC-D-279 EN.pdf - 第50页

−20°C to +20°C, in which the change from stress- to strain-driven solder response takes place, do not follow the damage mechanism described in Eqs. #1 and #2 [Ref. A-9: 34]. The damage mechanism is dif ferent than for mo…

pad centers, L

is sometimes referred to as the

distance from the neutral point (DNP);

= steady-state operating temperature for compo-

nent, substrate (T

for power dissipation in

component) during high temperature dwell;

C,0

S,O

= steady-state operating temperature for compo-

nent, substrate during low temperature dwell,

for non-operational (power off) half-cycles

C,O

S,0

;

= (1/4)(T

C,O

S,O

), mean cyclic solder

joint temperature;

= CTEs for component, substrate;

∆D = potential cyclic fatigue damage at complete

stress relaxation;

∆T

−T

C,O

, cyclic temperature swing for compo-

nent;

∆T

−T

S,O

, cycling temperature swing for sub-

strate (at component);

∆(α∆T) = ?α

∆T

−α

∆T

?, absolute cyclic expansion

mismatch, accounting for the effects of power

dissipation within the component as well as

temperature variations external to the compo-

nent;

∆α = ?α

−α

?, absolute difference in CTEs of com-

ponent and substrate, CTE-mismatch, because

of the inherent variability in material proper-

ties ∆α<2x10

-6

should not be used in calculat-

ing reliability.

A-3.3 CAVEAT 1 — Solder Joint Quality

The solder joint fatigue behavior and the resulting reliabil-

ity prediction equations. Eqs. A-1 through A-2, were deter-

mined from thermal cycling results of solder joints that

failed as a result of fracture of the solder, albeit sometimes

close to the IMC layers. For solder joints for which layered

structures are interposed between the base material and the

solder joints, these equations could be optimistic upper

bounds if the interposed layered structures become the

‘weakest link’ in the surface mount solder attachments.

Such layered structures could be: metallization layers that

have weak bonds to the underlying base material, or are

weak themselves, or dissolve essentially completely in the

solder; oxide or contamination layers preventing a proper

metallurgical bond of the solder to the underlying metal;

brittle IMC layers too thick due to too many or improperly

long high temperature processing steps.

Some material choices can lead to lower quality and

weaker solder joints because the material is more difficult

to wet and solder. The nickel/iron alloys, Kovar

and

Alloy 42, fall into this material category. The resulting

lower solder joint quality indicated in Table A-2 is also

clearly evident in Figure A-2, where the solder joint pull

strength is shown for a variety of differently prepared

Alloy 42 and copper leads. Alloy 42 leads, even when

etched or pre-ﬂowed at temperatures higher than can be

tolerated by the component, show a substantial reduction in

the solder joint pull strength relative to copper. In the worst

instance, the leads from one Alloy 42 manufacturer have a

pull strength of less than half of those with more typical

Alloy 42 and are essentially non-wettable.

Early failures of the solder attachments of components with

Alloy 42 lead frames and leads during accelerated testing

[Refs. A-9: 26-30] and manufacture [Ref. A-9: 31] have

been documented.

Solder joints which have solder joint heights (gaps) of

h<50 to 75µm also require special attention. For solder

joints that thin, the gap is essentially ﬁlled with intermetal-

lic compounds and those solder metals that do not go into

solution with the base metals to form the IMCs. Therefore

Eqs. A-1 and A-2 do not apply because these gaps are not

ﬁlled with solder [Ref. A-9: 32]. These materials do not

creep as readily, if at all, at the prevailing temperatures and

are typically more brittle, but much stronger than solder.

Thus, fatigue lives are longer than would be predicted from

Eqs. A-1 and A-2 unless overstress conditions occur.

On the other hand, the fatigue lives of solder attachments

can be underestimated by Eqs. A-1 through A-4 if the com-

ponent is underﬁlled with a load-bearing substance, e.g.,

epoxy [Ref. A-9: 33]. Components that are glued-down to

the substrate result in higher solder joint fatigue reliability,

since the solder joints are loaded in compression when the

adhesive contracts on cooling from the solder reﬂow tem-

peratures. Covercoats can either increase or decrease solder

joint fatigue lives depending on the properties of the cover-

coat and when and how it is applied. Parylene

has been

found to increase the solder joint fatigue life by about a

factor of three.

In general, caution might be indicated in all instances

where the predicted life is less than 1000 cycles, because

the severe loading conditions producing such short lives

are likely to produce different damage mechanisms or/and

failure modes.

A-3.4 CAVEAT 2 — Large Temperature Excursions

Solder joints experiencing large temperature swings

(−50°C to + 80°C) which extend both signiﬁcantly below

and signiﬁcantly above the temperature region bounded by

Table A−2 Quality of Solder Joints with Copper and

Alloy 42 Resulting from Different Reﬂow Temperatures

Reﬂow

Temperature

°C

Solder Joint Quality

60/40 Solder to

Copper

60/40 Solder to

Alloy 42

~210 just o.k. marginal to bad

~240 good just o.k.

~260 good good

IPC-D-279 July 1996

−20°C to +20°C, in which the change from stress- to

strain-driven solder response takes place, do not follow the

damage mechanism described in Eqs. #1 and #2 [Ref. A-9:

34]. The damage mechanism is different than for more

typical use conditions and is likely dependent on a combi-

nation of creep-fatigue, causing early micro-crack forma-

tion, and stress concentrations at these micro-cracks caus-

ing faster crack propagation during the high stress cold

temperature excursions, as well as recrystallisation consid-

erations.

A-3.5 CAVEAT 3 — High-Frequency/Low-Temperatures

For high-frequency applications, f>0.5 Hz or t

<l s, e.g.,

vibration, and/or low temperature applications, T

< 0°C,

for which the stress relaxation and creep in the solder joint

is not the dominant mechanism, the direct application of

the Coffin-Manson [Ref. A-9: 14] fatigue relationship

might be more appropriate. This relationship is

(50%)=

[

f’

∆γ

]

−1

[Eq. A-5]

where ∆γ

is the cyclic plastic strain range and c ≈ −0.6.

It has to be noted, that the determination of ∆γ

depends on

the expansion mismatch displacements and the separation

of the plastic from the elastic strains.

For loading conditions of this character, it is possible that

high-cycle fatigue behavior may be observed.

A-3.6 CAVEAT 4 — Local Expansion Mismatch

For applications for which the global thermal expansion

mismatch is very small, e.g. ceramic-on-ceramic or silicon-

on-silicon (ﬂip-chip solder joints), the local thermal expan-

sion mismatch becomes the primary cause of fatigue dam-

age. Equation A-4 does not address the local thermal

expansion mismatch. This reliability problem needs to be

assessed using an interfacial stress analysis [Ref. A-9: 35]

and appropriate accelerated testing.

For leaded surface mount components with lead materials

that have CTEs signiﬁcantly lower than copper alloy mate-

rials, e.g., Kovar

or Alloy 42, the results from Eqs. A-1

and A-2 will be optimistic, since the fatigue damage con-

tributions from the solder/lead material CTE-mismatch, the

local thermal expansion mismatch, are not included.

It has shown that the interfacial stresses resulting from the

local expansion mismatch follow [Ref. A-9: 35].

τ∝L (α

Solder

−α

Basr

)(T

max

− T

min

)

[Eq. A-6]

Figure A−2 Solder Joint Pull Strengths for Gullwing Leads Consisting of Alloy 42 from Different Vendors and Copper

[Ref. A-9: 31]

July 1996 IPC-D-279

where L is the wetted length of the solder joint. In addition,

besides substantial shear stresses at the interface between

the solder joint and the base material to which it is wetted,

even larger peeling stresses occur. Both of these stresses

are proportional to the parameters given in Eq. A-6.

From Eq. A-6 it is quite clear, that for leads consisting of

Alloy 42, the wetted length of the solder joint, that is the

length of the lead foot should be minimized to reduce inter-

facial stresses. That, of course, is contrary to the good

practice that the foot length should be at least three times

the lead width for optimum solder joint quality. However,

since in most applications, the local expansion mismatch

results in contributory damage to the more important dam-

age caused by the global expansion mismatch, this contra-

indication can be ignored without suffering catastrophic

consequences.

From the available experimental data, the damage term, to

be used in Eq. A-1, for the local expansion mismatch alone

∆D(local)=

[

L∆α∆T

]

[Eq. A-7]

where the parameters are the same as in Eq. A-6 and

=0.1 mm, a scaling wetted length. The local expansion

mismatch is then treated as an additional loading condition

(see sections A-3.9 & A-3.10).

A.3.7 CAVEAT 5 — Very Stiff Leads/Very Large Expan-

sion Mismatches

Equations A-3 and A-4 differentiate between surface mount

solder attachments that are leadless and those with compli-

ant leads. Leadless solder attachments presume substantial

plastic strains due to yielding prior to creep and stress

relaxation, whereas Eq. A-4 assumes that the compliant

leads prevent stresses in the solder joints to reach levels

where substantial yielding, and thus plastic strains prior to

creep and stress relaxation, can take place.

However, there is an intermediate region that is not cov-

ered by these assumptions. For very stiff, non-compliant

leads (e.g., SM connector headers), perhaps at lead stiff-

nesses K

> ~90 N/mm and/or for very large thermal

expansion mismatches (e.g., ceramic MCMs on FR-4)

resulting in strain ranges ∆γ > ~10%, the damage estimates

in Eq. A-4 can be substantially in error, because the

assumptions underlying Eq. A-4 are violated.

For very stiff leads the stresses calculated in Eq. A-4 can

exceed the yield strength of the solder. Since yielding will

not permit stresses signiﬁcantly higher than the yield

strength, these calculated stress ranges will overestimate

the cyclic fatigue damage and thus result in substantially

underpredicted fatigue lives. To prevent this analytical

error, the stress range in Eq. A-4 needs to be bounded by

the yield strength of solder in shear.

For very large thermal expansion mismatches the full dis-

placements will not be transmitted to the solder joints,

because the leads will accommodate displacements by plas-

tic deformations of the lead material. Possible exceptions

are situations where very stiff leads are also involved, in

which case the solder joint reliability is best estimated

using Eq. A-1 for leadless solder attachments. The strain

range that can be accommodated by creep and stress relax-

ation in the solder joints can be signiﬁcantly exceeded by

the displacements resulting from very large thermal expan-

sion mismatches and the cyclic fatigue damage would be

signiﬁcantly overestimated. Under these conditions FEA is

required to determine the split in the accommodation of the

displacements between the lead and the solder joints.

Under these circumstances, Eqs. A-3 and A-4 will provide

lower and upper bounds for the reliability estimates,

respectively. The higher the lead stiffness, the closer the

expected results will be towards the results given by Eq.

A-3 for the leadless—‘inﬁnitely stiff leads’—solder attach-

ments. Very high lead stiffnesses can occur in the case of

through-hole component leads converted to surface mount

and for connector headers where the male header pins have

been simply bent into a gull-wing lead foot without any

reduction in the lead cross-section. Very high thermal

expansion mismatches occur primarily in accelerated test-

ing and in extraordinary environments like storage and

transport for products that are designed for benign operat-

ing environments.

A.3.8 Statistical Failure Distribution and Failure Prob-

ability

While the physical parameters deﬁne the median cyclic

fatigue life from physics-of-failure considerations, solder

attachment failures for a group of identical components

will follow a distribution—like all fatigue results—which

typically is best described by a Weibull statistical distribu-

tion [Ref. A-9: 36]. Given the statistical distribution, the

fatigue life at any given failure probability for the solder

attachment of a component can be predicted as long as the

slope of the Weibull distribution is known. Thus, the

fatigue life of surface mount solder attachments at a given

acceptable cumulative failure probability per component, x,

is —assuming a two-parameter (2P) Weibull statistical

distribution—given by

(x%)=N

(50%)

[

1n(1 − 0.01x)

1n(0.5)

]

[Eq. A-8]

where β = Weibull shape parameter or slope of the

Weibull probability plot; typically β≈3 for

fatigue tests, from low-acceleration tests of

stiff leadless solder attachments β≈4 and ≈2

for compliant leaded attachments.

IPC-D-279 July 1996